FE
562
. K3
no.
r iWA
RD
80172
Report No. FHWA/RD80/172
IANDE1980: BOX CULVERTS AND SOIL MODELS
May 1981
Final Report
*Mm o*
Document is available to the public through
the National Technical Information Service,
Springfield, Virginia 22161
Prepared for
FEDERAL HIGHWAY ADMINISTRATION
Offices of Research & Development
Structures and Applied Mechanics Division
Washington, D.C. 20590
FOREWORD
This report presents the results of a study designed to extend the
capability of the FHWA "CANDE" (Culvert Analysis and Design) computer
program to include the capability for the automated finite element
analysis for the structural design of precast reinforced concrete box
culvert installations. The study also resulted in a new reinforced
concrete model with loading through ultimate, unloading and redistribution
of stresses due to cracking, as well as a new soil model (the socalled
Duncan model). Included in the report is a User Manual Supplement and
three (3) solved sample problems. Overlay instructions permit the
program to be executed more efficiently and with less computer core
storage requirements. This report will be of primary interest to
supervisors, engineers, and consultants responsible for the design of
culverts.
This report is being distributed under FHWA Bulletin with sufficient
copies of the report to provide one copy to each regional office, one
copy to each division, and one copy to each State highway department.
Direct distribution is being made to the division offices.
"Charles F. Scheffey
"j x Director, Office of Research
/ Federal Highway Administration
NOTICE
This document is disseminated under the sponsorship of the Department of
Transportation in the interest of information exchange. The United
States Government assumes no liability for its contents or use thereof.
The contents of this report reflect the views of the authors, who are
responsible for the facts and the accuracy of the data presented herein.
The contents do not necessarily reflect the official views or policy of
the Department of Transportation.
This report does not constitute a standard, specification, or regulation,
The United States Government does not endorse products or manufacturers.
Trademarks or manufacturers' names appear herein only because they are
considered essential to the object of this document.
Technical Report Documentation Page
1. Report No.
FHWA/RD80/172
2. Government Accession No.
3. Recipient's Catalog No.
4. Title and Subtitle
CANDE1980: Box Culverts and Soil Models
5. Report Date
May 1981
6. Performing Organization Code
7. Author's)
Katona, M.G, Vittes, P.P., Lee, C.H., and Ho, H.T.
8. Performing Organization Report No.
9. Performing Organization Name and Address
University of Notre Dame
Notre Dame, Indiana 46556
10. Work Unit No. (TRAIS)
3513241
11. Contract or Gront No.
D0TFH1 19408
12. Sponsoring Agency Name and Address
Offices of Research and Development
Federal Highway Administration
U.S. Department of Transportation
Washington, D.C. 20590
13. Type of Report and Period Covered
Final Report
November 1978October 1980
DEPARTMENT
TRANSPORTATION
\$}Y Spc isoring Agency Code
15. Supplementary Notes
George W. Ring, Contract Manager, HRS14
JAN11198Z
\—
LiDHAnV
16. Abstract
The CANDE computer program, introduced in 1 970 Fur lilt! iJTTUctijral design and analysis
of buried culverts, is extended and enhanced in this work effort to include options
for automated finite element analysis of precast, reinforced concrete box culverts,
and new nonlinear soil models. User input instructions for the new options, now
operative in the CANDE1980 program, are provided in the appendix of this report
along with example input/output data.
Comparisons between CANDE1980 predictions and the elastic analysis/design method
used to develop the ASTM C789 design tables for precast box culverts revealed the
importance of soil structure interaction which is not taken into account in the latter
method. As a general conclusion, the ASTM C 789 design tables provide safe designs
(conservative) providing that good quality soil is used for backfill.
The socalled Duncan soil model, employing hyperbolic functions for Young's modulus
and bulk modulus, is a new soil model option in CANDE1980. Standard soil model
parameters, established from a large data base of triaxial tests, are stored in the
program and can be used by simply identifying the type of soil and degree of compac
tion. In a similar manner, simplified data input options have also been developed
for the overburden dependent soil model.
In addition to user input instructions and example input/output data, the appendices
also provide overlay instructions to reduce computer core storage requirements.
17. Key Words
Culverts, Box Culverts, Soil Models,
Soil Structure Interaction
18. Distribution Statement
No restrictions. This document is available
through the National Technical Information
Service, Springfield, Virginia 22161.
19. Security Classif. (of this report)
Unclassified
20. Security Classif. (of this page)
Unclassified
21. No. of Pages
214
22. Price
Form DOT F 1700.7 (872)
Reproduction of completed page authorized
ACKNOWLEDGMENTS
Representatives of industry, state highway departments, universities
and research groups have been very helpful in providing information and
constructive comments for this research effort. A special thank you is
extended to Dr. James M. Duncan of the University of California for
providing data and details of his soil model, and to Dr. Frank J. Heger
of Simpson Gumpertz and Heger Inc. who, along with representatives of
the American Concrete Pipe Association, supplied experimental data for
outofground tests. Mr. Robert Thacker provided consultation on over
laying the CANDE program on the IBM computers and programming the metric
version of CANDE1980.
TABLE OF CONTENTS
p age
CHAPTER 1  INTRODUCTION 1
1.1 Background 1
1.2 Objectives 2
1.3 Scope and Approach 2
CHAPTER 2  REVIEW OF PRECAST BOX CULVERTS 4
2.1 Background 4
2.2 Development and rational of ASTM precast
box culvert standards 5
CHAPTER 3  REINFORCED CONCRETE MODEL 10
3.1 Objective 10
3.2 Assumptions, limitations and approach 10
3.3 Basic formulation for beamrod element 12
3.4 Finite element interpolations 16
3.5 Stressstrain relationships 19
3.6 Section properties 23
3.7 Incremental solution strategy 24
3.8 Measures of reinforced concrete performance 26
3.9 Standard parameters for concrete and reinforcement .... 29
CHAPTER 4  EVALUATION OF REINFORCED CONCRETE MODEL FOR
CIRCULAR PIPE LOADED IN THREEEDGE BEARING 31
4.1 Preliminary investigations 31
4.2 Experimental tests 33
4.3 Analytical model and comparison of results 37
CHAPTER 5  EVALUATION OF REINFORCED CONCRETE MODEL FOR
BOX CULVERTS LOADED IN FOUREDGE BEARING 49
5.1 Experimental tests 49
5.2 CANDE model 53
5.3 Comparison of models with experiments 53
53
CHAPTER 6  DEVELOPMENT OF LEVEL 2 BOX MESH 62
6.1 Parameters to define the models 62
6.2 Assumptions and limitations 67
CHAPTER 7  EVALUATION OF CANDE BOXSOIL SYSTEM 71
7.1 Sensitivity of soil parameters 71
7.2 Comparison with test data  79
in
TABLE OF CONTENTS (Continued)
Page
CHAPTER 8  EVALUATION OF ASTM C789 DESIGN TABLES WITH CANDE 85
8.1 Box section studies for dead load 85
8.2 Box section studies with live loads Ill
CHAPTER 9  SOIL MODELS 115
9.1 Duncan model representation of elastic parameters 117
9.2 CANDE solution strategy for Duncan model 121
9.3 Standard hyperbolic parameters 125
CHAPTER 10  SUMMARY AND CONCLUSIONS 132
APPENDICES
A  Details of reinforced concrete model 134
B  CANDE1980; User Manual Supplement 147
C  Sample of input data and output 177
D  System overlay 202
REFERENCES 208
IV
CHAPTER 1
INTRODUCTION
1.1 BACKGROUND
The CANDE computer program (Culvert ANalysis and Design) was first
introduced in 1976 for the structural analysis and design of buried cul
verts (1,2,3). CANDE employes soilstructure interaction analysis and
has a variety of options, such as; choice of culvert type (corrugated
steel, corrugated aluminum, reinforced concrete, and plastic) and choice
of analysis /design method (elasticity solution  level 1, automated finite
element solution  level 2, and standard finite element solution  level 3)
Other features include; linear and nonlinear culvert and soil models, in
cremental construction and soilstructure interface elements.
Since its introduction in 1976, the program has been widely distri
buted and used by state highway departments, federal agencies, consulting
firms, industry, research laboratories, and universities in the United
States and Canada. User responses have been very favorable along with
encouragement and suggestions for extending the program's capabilities.
In particular, it is observed that reinforced concrete box culverts have
dramatically increased in use during recent years. To analyze these with
CANDE (1976 version) requires level 3 analysis with time consuming finite
element data preparation. Prior to this work, the automated finite element
level 2 analysis was restricted to round or elliptical pipes. Thus, a
desirable program extension is a level 2 analysis for box culverts with
the capability to analyze through ultimate loading. A second observation
is the wide spread popularity of the socalled Duncan soil model (26, 27,
28, 29) which has been developing over the last decade and is formulated
on a large experimental data base for many types of soil. The above
observations lead to the objectives of this work.
1.2 OBJECTIVES
The first major objective is to develop and incorporate into the
CANDE program an automated finite element analysis solution method for
buried, precast reinforced concrete box culverts, called here, "level 2
box" option. Included in this objective is validating the CANDE model
with experimental data for loadings through ultimate and comparisons
with other design/analysis methods.
The second major objective is to incorporate the Duncan soil model
into the CANDE program with due regard to convergence problems and to
provide options for simplified data input.
1.3 SCOPE AND APPROACH
To meet the above objectives, a step by step approach was undertaken
for both major goals. First, for the development and validation of pre
cast reinforced concrete box culverts, the steps are:
(a) Review current design/analysis procedures to assess the state
oftheart and to establish a comparitive basis with CANDE
(Chapter 2).
(b) Reformulate the existing reinforced concrete model to include
loading through ultimate, unloading, and redistribution of
stresses due to concrete cracking (Chapter 3 and Appendix A).
(c) Evaluate and validate the reinforced concrete model with out
ofground experimental data including pipes with 3edge bearing
loads and boxes with 4edge bearing loads (Chapters 4 and 5).
(d) Develop an automated finite element solution method (level 2
box) for buried box culverts (boxsoil model) with simplified
input for embankment and trench installations (Chapter 6).
(e) Evaluate and validate the boxsoil model with available experi
mental data and parametric studies (Chapter 7).
(f) Cross check the boxsoil model predictions with current design/
analysis procedures in step (a) and evaluate current design
methods (Chapter 8).
Next, for the objective of incorporating the Duncan soil model and
simplifying soil model input, the steps are (Chapter 9):
(a) Evaluate the Duncan soil model to verify reasonable behavior
in confined compression and triaxial loading.
(b) Investigate iterative solution strategies to enhance convergence
and incorporate the model into CANDE program.
(c) Establish standard model parameters dependent on soil type and
degree of compaction for the simplified data input option.
Also, simplify data input for the existing overburden dependent
soil model.
All program modifications noted above have been incorporated into
CANDE, hereafter called CANDE1980 to distinguish it from the 1976 version.
Appendix B provides input instructions to exercise the new options contained
in CANDE1980. These instructions are a supplement to the 1976 CANDE User
Manual (2) and only need to be referred to if the new options are desired.
In other words, the 1976 user manual is compatible with the CANDE1980
program. Appendix C illustrates inputoutput data for some of the new
options and Appendix D provides system overlay instructions to reduce
core storage.
The CANDE1980 program discussed herein is based on the English
system of units. A companion program in metric units has been developed
and is also available from FHWA.
CHAPTER 2
REVIEW OF PRECAST BOX CULVERTS
In this chapter a brief review on the development of precast rein
forced concrete box culverts is presented along with a discussion of
current design procedures. The intent is to acquaint the reader with
precast box culverts, terminology and design concepts and to "set the
stage" for the CANDE methodology presented in later chapters. For
brevity, "reinforced concrete box culverts" will be referred to as "box
culverts".
2.1 BACKGROUND
Precast box culverts, as opposed to castinplace box culverts, are
relatively recent additions in culvert technology, coming into popular
use within the last decade. For many years, castinplace box culverts
have been used in installations with special requirements or by design
preference. However, castinplace culverts have inherent disadvantages;
high labor costs associated with castinplace construction, lengthy
periods of traffic disruption, and minimal quality control often compensated
for by conservative designs. Alternatively, plantproduced box culverts,
manufactured under strict quality control and installed by rapid cutand
fill procedures, can offset these disadvantages particularly if the box
dimensions, reinforcement, ect., are standardized for manufacture.
With the above motivation, the Virginia Department of Highways
and the American Concrete Pipe Association (ACPA) , with financial support
from the Wire Reinforcement Institute, initiated a cooperative program,
early in 1971, to develop manufacturing specifications and standard
designs for precast box culverts. These specifications were to be
adaptable as a national standard under the auspices of the American
Society of Testing Materials (ASTM) and the American Association of State
Highway and Transportation Officials (AASHTO) . To this end, ACPA contracted
the consulting firm of Simpson, Gumpertz and Heger Inc. (SGH) to develop
a computerized design program for precast box culverts in cooperation
with ASTM committee C13.
Ultimately, this effort culminated in the ASTM C789 and AASHTO M259
specifications on Precast Reinforced Concrete Box Sections for Culverts,
Storm Drains and Sewers first published in 1974. These specifications were
limited to box culverts with a minimum of two feet (0.61 m) of earth cover.
Further developmental work by SGH resulted in the additional specifications
ASTM C850 and AASHTO M273 published in 1976 for precast box culvert in
stallations with less than two feet of earth cover. The above ASTM and
AASHTO specifications are essentially the same except for a few details
which are apparently now resolved. For purposes of this study, the
ASTM specifications will be used as reference. Design methods for pre
cast box culverts, other than those embodied in ASTM or AASHTO specifi
cations, will not be reviewed here since they are not standardized nor
have they gained national acceptance. Recently, ACPA published a sur
vey (Concrete Pipe News, June 1980) showing that usage of precast box
culverts, designed by ASTM specifications, has increased dramatically
within the last year. . . the number of projects and linear footage
installed in 1979 is almost equal to the total for the previous five years!
2.2 DEVELOPMENT AND RATIONALE OF ASTM PRECAST BOX CULVERT STANDARDS
The SGH computerized design program (12) is the basis of the
design rationale in the ASTM C789 design tables. A typical box cross
section is shown in Figure 2.1 along with nomenclature. The SGH design/
analysis approach includes the following steps; (a) load distributions
are assumed around the culvert in an attempt to simulate dead earth loads
and live loads, (b) moment, shear, and thrust distributions are determined
by standard matrix methods using elastic, uncracked concrete section pro
perties, (c) in the design mode, steel areas are determined by an ulti
mate strength theory for bending and thrust, where ultimate moments and
thrusts are obtained from step (b) multiplied by a load factor, (d) crack
width (0.01 inch allowable) is checked using a semiempirical formula
M
^1
f
^
1
/\
T 1
4>
v At2 \
»
<M
R
:A s4
1
_,
k s1
<C
t.
T
A s i = outer reinforcement
A S 2= top inner reinforcement
A S 3= bottom inner reinforcement
A S 4= side inner reinforcement
C = cover distance of reinforcement, uniform
H=haunch dimension
M=minimum length A s1 steei, top and bottom
R=rise distance , inside to inside
S = span distance, inside to inside
T=wall thickness, uniform
Figure 2.1. Typical Box Culvert Cross Section
controlled by steel stress at service loads, (e) ultimate shear stress
( 2 /P~ ) is checked against the nominal shear stress obtained in step (b)
multipled by a load factor.
For the standard box sizes shown in Table 2.1, the SGH design program
was used to generate the ASTM C789 design tables wherein steel reinforce
ment requirements are specified as a function of design earth cover be
ginning with a two foot minimum.
In a similar manner, ASTM C850 design tables were generated for earth
covers less than two feet. Here, the SGH design procedure was modified
to include requirements for longitudinal steel design due to concentrated
live loads (see ASTM Symposium STP 630).
Although the SGH design/analysis program has not been validated
with experimental data from buried box culverts, fairly good correlation
with outofground experimental tests has been reported (13). More will
be said about these experiments in Chapter 5.
Experimental data for instrumented, buried box culverts is extremely
limited. As of this writing, only two state highway departments (Kentucky
and Illinois) are known to have undertaken experimental programs for in
strumenting (settlement, soil pressure, and strain gages) buried box
culvert installations. Other states have made visual inspection reports
on the performance of buried box installations, but this data has marginal
value for validating design/analysis procedures. Data from the Kentucky
Department of Transportation was made available for this study and is
used to evaluate the CANDE program in Chapter 7.
In summary, the ASTM design tables for buried, precast box culverts,
which are based on the SGH design/analysis program, have not been pre
viously validated with experimental data from buried installations. Nor
have the tables been crosschecked with analytical procedures, such as
CANDE, employing soilstructure interaction and the nonlinear nature of
reinforced concrete. With this goal in mind, a step by step approach is
presented in the following chapters. First, the theory of CANDE 's non
linear, reinforced concrete model is developed. Second, the model is
TABLE 2.1 Standard box sizes, ASTM C789
Span
ft.
2
3
4
5
Rise, ft.
6 7 8
9
10
Wall
Thickness
in.
3
X
X
4
4
X
X
X
5
5
X
X
X
6
6
X
X
X
X
7
7
X
X
X X
8
8
X
X
XXX
8
9
X
XXX
X
9
10
X
XXX
X
X
10
1 ft = 0.3048 m
1 in = 2.54 cm
validated with experimental data for outofground conditions. Third,
the reinforced concrete model is combined with soil system models and
compared with experimental data from a buried installation. Last, the
CANDE model is used to evaluate the ASTM design tables.
CHAPTER 3
REINFORCED CONCRETE MODEL
3.1 OBJECTIVE
A reinforced concrete, beamrod member, whether it be part of a
culvert or any other structural system, poses a difficult analysis
problem due to the nonlinear material behavior of concrete in com
pression, cracking of concrete in tension, yielding of reinforcement
steel, and the composite interaction of concrete and reinforcement.
Matters are further complicated when the internal loading is not
proportional, i.e., when the internal moment, shear and thrust at
a particular cross section change in different proportions (including
load reversals) during the loading history. Such is the case for
buried culverts during the installation process.
In this chapter, the development of a reinforced concrete beam
rod element is developed in the context of a finite element formulation
for CANDE1980. This model is more general than the model in CANDE1976
and includes; incremental loading through ultimate, unloading, and
redistribution of stresses due to cracking.
The following presentation provides an overview of the model
development emphasizing assumptions and limitations. Details of the
numerical solution strategy are presented in Appendix A. Evaluation
of the model with experimental data and other theories is presented
in subsequent chapters.
3.2 ASSUMPTIONS, LIMITATIONS, AND APPROACH
Listed below are the fundamental assumptions for the reinforced
concrete beamrod element.
10
I
1. Geometry and loading conform to plane strain implying
the beam rod element is of unit width. Constant section
properties are assumed through an element length, but
may differ between elements.
2. Displacements and strains are small. No buckling consid
erations are included.
3. Planes remain plane in bending and shear deformation is
negligible.
4. Concrete is linear in tension up to cracking. Cracked
concrete cannot carry tension stresses and precrack
stresses are redistributed. In compression, concrete is
modeled with a trilinear stressstrain curve terminating
at ultimate strain. Unloading is elastic.
5. Reinforcement steel is elasticplastic and identical in
compression and tension. Unloading is elastic.
6. Reinforcement steel is lumped into two discrete points
near the top and bottom of the crosssection and deforms
with the crosssection,
7. Element lengths are sufficiently small so that the current
stress distribution through a crosssection is representative
of the entire element for purposes of computing current
section properties.
8. Loads are applied incrementally and sufficiently small
so that the stressstrain relations (for both steel and
concrete) can be regarded as incremental tangent relations
determined iteratively over the load step.
11
In overall perspective, the developmental steps begin with an
incremental statement of virtual work wherein the beamrod assumptions
are introduced along with standard finite element interpolation func
tions for axial and bending deflections. This results in a tangent
element stiffness matrix and incremental load vector that can be
assembled into a global set of system equations with unknown nodal
degrees of freedom, and solved by standard techniques (1). However,
the global matrix contains estimates of the bending and axial stiff
ness for each beamrod element (as well as estimates for soil stiff
ness if nonlinear soil models are part of the system). Thus, each
load step is repetitively solved (iterated), and the results are used
to improve the stiffness estimates until convergence is achieved.
Prior to the first loading increment, the beamrod element is
assumed stress free and uncracked so initial stiffnesses correspond
to an uncracked, elastic, transformed reinforced concrete cross
section. Upon applying the first load increment, the first tentative
solution may indicate that some elements should have had reduced
stiffnesses due to cracking or yielding of the section. Using the
strain distribution at the beginning and end of the load step, new
stiffness estimates are obtained and the process is iterated to con
vergence. Each subsequent load step is treated in a similar fashion
where a history of maximum stress and strain is maintained for pur
poses of identifying unloading conditions.
The above assumptions and general approach are outlined in the
following development,
3,3 BASIC FORMULATION FOR BEAM ROD ELEMENT
In this section we consider an incremental virtual work statement
for a unit width, beamrod element with body forces given by:
6AV = 5AU  6AW (3,1)
w ith 5AU = / / SeAa dxdy = internal virtual work increment
x y
12
SAW = f f 6{. 7 } { , 1} dxdy = external virtual work increment
x y v Af„
where a = normal stress, xdirection
e = normal strain, xdirection
u = longitudinal displacement, xdirection
v = transverse displacement, ydirection
f. = longitudinal body force, xdirection
f_ = transverse body force, ydirection
x = space coordinate parallel to beam axis
y = space coordinate transverse to beam axis
6 = virtual symbol
A = increment symbol
The above beam displacements are illustrated in Figure 3.1.
Introducing BernouliEuler beam kinematics (Assumptions 2 and 3),
longitudinal displacements through a cross section may be arbitrarily
decomposed into a uniform axial distribution, u (x), plus a distri
bution proportional to slope, v'(x), and linearly varying about some
axis y, i.e. :
u(x,y) = u Q (x) + v f (x) (yy) 3.2
Later, when the above kinematic relation is incorporated into Equation
3,1, the axis y will be chosen such that internal bending work is un
coupled from internal axial work,
Employing the small straindisplacement assumption, normal strain
is:
e(x,y) = u^(x) + v"(x) (5y) 3.3
where primes denote derivatives with respect to the argument.
To complete the field variable assumptions, a general, nonlinear
stressstrain relationship is assumed in incremental form as:
Aa = E'(e)Ae 3,4
13
4*
Figure 3.1 Deformation of BeamRod Element
AF
A t 1 MHt t f t f t t t
Ui /^R~ AF1
A
i>r«i
£>*
e
i
Figure 3.2 Nodal Degrees of Freedom
and Element Loading
14
Here E f (e) is a tangent modulus relating increments of stress to
increments of strain and is dependent on loading history. Naturally,
the functional forms of E'(e) are different for concrete and steel
materials. However for clarity of presentation, the specific forms of
E'(e) will be deferred to a later section.
Using the incremental form of Equation 3.3 along with Equation 3.4
and integrating through the cross section, the internal virtual work
increment may be expressed as:
5AU = / (6ui EA* Aul + 6v M EI* v" + EX*(5v"Au: + 6ulAv"))dx 3.5
x
*
where EA = / E' (e) dy = effective axial stiffness 3.6
y
* _ 2
EI = / E'(e) (yy) dy = effective bending
y stiffness 3.7
*
EX = / E'(e) (yy) dy = axialbending coupling 3.8
_ *
The location of y is now chosen so that the coupling term EX is
zero. Thus, y is given by:
y = ( / E»(e) y dy)/EA* 3.9
This choice of y is convenient because bending and axial deformations
are uncoupled in the virtual work statement. However, it must be
remembered that y, like EA and EI , is dependent on E'(e), thus these
values change during each load step.
To complete the virtual work statement, the kinematic assumption
(Equation 3.2) is introduced into the external virtual work incremen
tal expression and integrated over the cross section to give:
6AW = / ( 6u AF l + 6vAF 2 + <Sv?AF 3 ) dx 3 ' 10
where AF = / Af, dy = axial body force per unit length
1 y 1
AF„ = J" Af dy = transverse bodv force per unit length
2 y 2 J f
15
AF = / Af (yy) dy = body moment per unit length
■j y x
The body moment, AF , is generally nonzero except if the centroid of
the axial body weight happens to coincide with the current location
of y. However, the magnitude of the body moment is usually negligible
compared to the magnitude of internal moments which arise from trans
verse loading in culvert installations. Thus, the body moment is
neglected in this study.
Equations 3.5 and 3.10 are the internal and external virtual work
expressions for the beam rod element with unknown displacement functions
u (x) and v(x).
3.4 FINITE ELEMENT INTERPOLATIONS
Figure 3.2 shows a beamrod element with three nodal degrees of
freedom at each end node, an axial displacement, a vertical displace
ment, and a rotation. These degrees of freedom are used to define
admissible interpolation functions for u n (x) and v(x) in the context
of a finite element formulation.
Specifically, the axially displacement, u (x) is approximated
with a twopoint Lagrange interpolation function:
■■ * D]
u Q (x) = O, cj) > 3.11
where u, = axial displacement at node 1
u~ = axial displacement at node 2
(J> 1 (x) =16
4> 2 (x) = 3
B(X) = x/£
For transverse displacements, v(x), a twopoint Hermetian inter
polation function is used.
16
v(x) =
<a l a 2 a 3 a ?
v.
3.12
where
v = transverse displacement at node 1
v„ = transverse displacement at node 2
6.. = rotation at node 1
6„ = rotation at node 2
a 1 (x)
2 3
1  33 + 23
a 2 (x)
a 3 (x)
= 3(13) I
2 3
33  23
a 4 (x)
= 3 (3D I
Upon substituting the interpolation functions into the Incremental
virtual work expression, 6AV = 5AU6AW, we have:
e
6AV = <6r > {[K ] {Ar}  {AP }}
e e e
3.13
where
{r}
~* A
*■■
u l
*
v 1
J_
A
6 1
/v
U 2
*
9 2
/v
_ v 2
= element degrees of freedom 3.14
17
{AP } = jV
e 12
6AF ]
6AF,
JIAF,
i
6AF ]
6AF,
£AF,
= element load vector
3.15
[K ] =
e
* *
EA n n EA
\J
£
■k
12EI
6EI
* 3
* 2
*
«.
c.
*
EA
Symmetric.
(tangent element
stiffness matrix)
12EI
6EI
6EI
2EI
* *
12EI 6EI
4EI
The above tangent element stiffness and load vector are valid for
the local beam coordinates. For assembling element contributions into
the global coordinate system, standard coordinate transformation are
employed.
Note that the tangent element stiffness matrix is identical in
form to that obtained from standard matrix methods of structural anal
ysis. However, the axial stiffness EA and bending stiffness EI
(dependent on y) are not constant and must be determined iteratively
for each load step in accordance with Equations 3.6, 3.7, and 3.9.
These equations are dependent on the concrete and steel stressstrain
relationships discussed next.
18
3.5 STRESSSTRAIN RELATIONSHIPS
Concrete . The assumed stressstrain behavior for concrete is shown
in Figure 3.3 where the trilinear curve is defined by the following
input variables:
e = concrete strain at initial tensile cracking
e = concrete strain at initial elastic limit
y
e' = concrete strain at onset of ultimate
c
f ' = unconfined compressive strength of concrete
E = Young's modulus in linear zone
With the above input variables, three additional parameters can be
derived:
E n = (f  E e )/(e'  e ) = Young's modulus in yielding
2 c 1 y c y
J zone
£'  E, e = initial tensile strength
tit
f = E,e = initial yield strength
yc 1 y J b
In tension the concrete is linear until the initial tensile strain
exceeds the cracking strain limit e . When cracking occurs, the
tensile stress becomes abruptly zero (redistributed to noncracked
portions). Once a point in the cross section is cracked, the crack
does not heal, implying no tensile strength. Thus e is set to
zero for all subsequent reloading in tension.
For initial compression loading, the concrete begins to yield
with hardening at stress f . Perfect plasticity occurs at stress f^
and continues through ultimate strain. Unloading is elastic and
results in permanent plastic strains as indicated in Figure 3.4.
Reloading is elastic until the stress reaches its previous maximum
value after which it follows the original stressstrain curve. (See
Figure 3.4).
19
CO
(0
Cracking
> Strain
Figure 3.3 Idealized StressStrain Diagram for Concrete
unloadreload paths
i/ iE i A
/ /
/
/
^Strain
Figure 3.4 Elastic UnloadReload for Concrete
20
With the above understanding, the tangent modulus relationship
for concrete confined in a plane is expressed as:
E'(e) = E (1  a(e)) 3.17
c c
2
where E = E / (1  v )
c 1 c
with E = elastic, confined plane modulus of concrete
v ■ Poisson's ratio of concrete (constant)
c
a(e) = dimensionless function of stressstrain history
The dimensionless function a(e) ranges in value from 0.0 (elastic
response) to 1.0 (perfectlyplastic response), representing the non
linear effect of concrete. The actual value of a(e) to be used for
any given load increment is dependent on; known values of stress and
strain at the beginning of the step, known history parameters for
cracking and yielding, and unknown values of stress and strain at the
end of the step (iteration). Appendix A provides the details for
determining a(e) for all loading histories.
Steel . The assumed stressstrain behavior for reinforcing steel
is shown in Figure 3.5 where the elasticplastic curve is characterized
with two input variables:
E n = Young's modulus for steel
f = steel yield strength
Behavior in compression and tension is identical so that material is
elastic whenever the stress magnitude is less than f . Nonhardening
plastic flow occurs when the stress is equal to f , Unloading from
the plastic range is elastic and results in permanent plastic strains
(see Figure 3,5).
Similar to Equation 3,17 for concrete, the tangent modulus relation
ship for reinforcement steel confined in a plane is expressed as:
21
(0
©
CO
fy~
s
s
fy
^Strain
Figure 3.5 Idealized StressStrain Diagram
for Reinforcing Steel
X
unit I
width 7 '
Figure 3.6 Reinforced Concrete Cross Section
22
E' (e) = E (1  a(e)) 3.18
s s
2
where E = E J (1  v )
s s
with E = elastic, confined plane modulus of steel
s
v = Poisson's ratio of steel (constant)
s
a(e) ■ dimensionless function of stressstrain history
As in the case of concrete, the function a(e) for steel ranges in
value from 0,0 (elastic) to 1.0 (perfectly plastic) depending on
stressstrain history and stress values at the beginning and end of
each load step (see Appendix A).
3.6 SECTION PROPERTIES
Equations 3,17 and 3.18 represent the tangent modulus relation
ships for concrete and steel, respectively, which now can be used
to evaluate current section properties EA , y, and EI ,
Referring to a typical cross section shown in Figure 3.6, the
effective axial stiffness (Equation 3.6), the bending axis (Equation
3.9), and the effective bending stiffness (Equation 3.7) can be ev
aluated by separating the concrete and steel integration areas as
shown below.
* A
EA =
E'(e) dy + A . E» (e . ) + A _ E» (e A ) 3.19
c si s i sO s
y = <[ E»( £ ) ydy + A . E' (e.) y. + A . E' (e )y )/EA* 3.20
J j c v ' J si s i J x sO s
EI =
rh 2 _.,., ,2
E* (?) (yy) dy + A . E' (e . ) (yy.)
c si si i
+ a so e ; < i o> (?  y o )2 3  21
23
where A . = bottom steel area per unit width
. si
A  = top steel area per unit width
sO
y. = distance to A . from bottom
y i sx
y^ = distance to A _ from bottom
J sO
The integrals containing E'(e) represent the concrete contribution
c
to section properties and are evaluated numerically with 11point
Simpson integration, A stressstrain history is maintained at each
integration point for determining the current values of a(e). Steel
contributions to section properties are governed by E' (e ) and E'(e )
S X s u
representing the tangent steel modulus at the centroid of bottom and
top steel reinforcement.
The above equations suggest that the concrete contributions are
integrated over the entire section area irrespective of "holes" where
steel exists, however, the algorithm used in this study accounts for
these holes. These and other details of computing section properties
are discussed in Appendix A,
3.7 INCREMENTAL SOLUTION STRATEGY
All the assumptions and derivations for the beamrod element
have been presented. An overview of the solution strategy is given
next.
It is assumed that a converged solution is known at load step
i1 and it is desired to obtain a converged solution at load step i.
Basically, the objective is to determine effective section properties,
*  *
EI , y, and EI for each beam rod element.
A flow chart of the solution strategy is illustrated in Figure. 3. 7.
The procedure begins by initially assuming the section properties are
the same as the previous load step. Next, the system is assembled
for the current load increment and trial solutions are obtained for
moment and thrust increments in each element, given by:
24
* A
Estimate EA , y, EI for each element from
load step i1.
Apply load increment and solve system.
Obtain trial moment, AM, and thrust, AN,
increments for load step i.
3.
Estimate new strain distribution at load
step i as:
AN
= e ■ i + — *
ll *
EA
+ — * (yy)
EI
No
No
5.
* _
4. Compute new estimates for EA , y, and EI
(Equations 3.19, 3.20, and 3.21).
Test for inner loop convergence, i...e._, ,
Are two successive estimates of EA , y, and EI
(computed in Step 4) equal?
yes
6. Compute moment and thrust that must be
redistributed due to cracking.
Test for outer loop convergence, i.e., Are two
successive estimates of EA , y, and EI used in
Step 2 equal? If not, return to Step 2 and in Step 3
add the effects of redistribution (first time only).
yes
8. Converged solution increment. Sum incremental
responses to total response. Advance the load
step (i > i+1) and return to Step 1.
Figure 3.7. Flow chart of solution strategy.
25
AN = /. Aa dA = EA Au'
A.
AM  f Aa (yy) dA  EI Av"
A
Using the above relations together with Equation 3.3, a new strain
distribution is estimated as shown in Step 3 of the flow chart. This,
in turn, permits improving the estimates for section properties in the
"inner loop" iteration; steps 3, 4, and 5. Here, AM and AN remain
fixed (as estimated in Step 2) while the corresponding section
properties are determined. Note that inner loop operations are at
the element level, requiring no global assembly or solution.
Each time the inner loop converges, the converged section proper
ties are used in Step 2 to get new global solutions for AM and AN.
This process is called "outer loop" iteration and continues until two
successive solutions are equal within a specified tolerance. When
this occurs, convergence is achieved and the program advances to the
next load step (see Appendix A for additional detail),
3.8 MEASURES OF REINFORCED CONCRETE PERFORMANCE
Once a converged solution is obtained, measures of structural
distress are assessed by; (a) maximum tensile stress in steel (b),
maximum compressive stress in concrete, (c) maximum shear stress in
concrete, and (d) maximum crackwidth in concrete. The first three
measures of distress are evaluated directly from the structural
response predictions from the CANDE model, however the crackwidth
prediction employs a semiempirical approach. Each distress measure
is normalized by a corresponding design criterion to produce perform
ance factors as discussed below.
Steel Tension . The performance factor for steel reinforcing is
given by:
PF , = f /f
steel y max
where f = maximum steel stress (predicted)
max r
f = steel yield stress
y
26
For properly designed structures, this performance factor should be
in the range of 1,5 to 2,0, When the steel begins to yield, the
performance factor becomes 1,0 and remains .there through ultimate
loading.
Concrete Compression . For the outer concrete fibers experiencing
compressive stress from thrust and bending, the performance factor is:
PF = V/o
comp . c max
where a = maximum compressive stress (predicted)
f* = compressive strength of concrete
c
Proper designs should have this performance factor in the range 1.6 to
2.5, The performance factor remains at 1.0 when the concrete becomes
perfectly plastic and remains there through ultimate loading.
Concrete Shear . Nominal shear stress through a cross section is used
to define the shear performance factor, given by:
PF . o v /v
shear c max
where v = nominal average shear stress on section
max °
v = nominal concrete shear strength
c
Here v is computed by dividing the maximum predicted shear force by
the concrete area minus the cover area of steel. This definition is
consistent with the standard ACI measure of shear strength for beams
given by:
v  2.0 TV (psi)
c c
Other measures of shear strength are examined in the next chapter
with experimental data,
For proper design, the above performance factor should be in
the range 1.7 to 2.7. In the absence of stirrups, shear failure
(e.g. diagonal cracking) is assumed to occur when the performance
27
factor value is 1.0. Note that the CANDE model does not incorporate
diagonal cracking into the stressstrain law, only flexural cracking.
Concrete Crackwidth . The crackwidth prediction, C , is a semi
empirical approach wherein the maximum tensile steel stress predicted
by CANDE is used in an empirical formula proposed by Gergely and Lutz
(10). Using 0.01 inches (0.0254 cm) as the design standard for allowable
crackwidth, the cracking, performance factor is defined as:
PP , » 0.01/C
crack w
3
/£?
where C  Q.091/2t* S (f  5000)R (inches)
W D S
R ■ 1,34 x 10 (dimensionless number for culvert slabs).
t. ■ concrete cover to steel centroid (inches)
D
f ■ tensile steel stress (psi)
S ■ spacing of reinforcement (inches)
The Gergely and Lutz formula for C was found to give good pre
w
dictions for crackwidths in this study. This finding is further
supported by Lloyd, Relaji and Kesler (11) in their experimental tests
on oneway slabs with deformed wire, deformed wire fabric, and deformed
bars. The new crackwidth formula defined above replaces the old
crackwidth formula in CANDE 19 76. The new formula can be made to be
2
identical to the old by defining S = 0.68/A t^ where A is tension
2 sos
steel reinforcement, in /in.
Ultimate Loads . Ultimate loading in thrust and bending occurs in
a .beam rod element when the reinforced concrete section cannot sustain
any additional loading, i.e., all uncracked concrete is at maximum
compressive strength f ' and all reinforcement steel is yielding (plastic
hinging). For a structure composed of beamrod elements, such as a
box culvert, ultimate loading occurs when a sufficient number of plastic
hinges have formed to produce a collapse mechanism. This can be de
termined from the CANDE program by observing unrestrained deformation
as the load is increased to ultimate.
28
Ultimate loading in shear is assumed to occur when the performance
factor for shear in any beamrod element becomes 1,0. If a structure
fails in shear prior to flexuralthrust failures, the CANDE model is
still capable of carrying load up to flexuralthrust failure because
diagonal cracking is not included in the model development. Thus for
loads exceeding concrete shear failure, it must be presumed that suf
ficient shear reinforcement (stirrups) is available.
3.9 STANDARD PARAMETERS FOR CONCRETE AND REINFORCEMENT
Based on investigations presented in subsequent chapters, a set
of standard parameter values for concrete is given in Table 3,1 (see
also Figure 3.3). Except for compressive strength f and cracking
strain e , the parameters are assigned unique values, some of which
are dependent on f*.'
For subsequent analytical studies, the concrete will be
characterized by specifying f and £ . The remaining parameters
c t
are assigned the standard values shown in Table 3.1 unless stated
otherwise.
Standard parameters for reinforcement steel are shown in Table 3.2
wherein the yield stress in considered as the primary variable.
29
Table 3.1
Standard Concrete Parameters
Parameter
Symbol
Value
Compressive strength
Elastic modulus
Cracking Strain
Initial yield strain
Strain at f
c
Weight density
Poisson's ratio
f
c
3000 to 7000 (psi)
33/F (y ) 1,5 (psi)
c c
0,0 to 0,0001
(in/in)
0.5 f /E,
c 1
(in/ in)
0,002
(in/in)
150
(lbs/ft 3 )
0.17
_
Table 3.2
Standard Steel Parameters
Parameter
Symbol
Value
Yield strengh
Elastic modulus
Poisson's ratio
30 to 90
ksi
29000
ksi
0.3
_
1 psi = 6.895 kPa
1 pcf = 157.1 N/m~
30
CHAPTER 4
EVALUATION OF REINFORCED CONCRETE MODEL
FOR CIRCULAR PIPE LOADED IN THREEEDGE BEARING
In this chapter the validity of the reinforced concrete model
(presented in the previous chapter) is examined by comparing results
with experimental data for circular pipe tested outofground in three
edge bearing, i.e., the socalled Dload test (ASTM C49765T), The
objective is to determine if the model can reasonably predict load
deflection histories, the load at which 0.01 inch (0.254 cm) crackwidths
occur, and ultimate load.
4.1 PRELIMINARY INVESTIGATIONS
Prior to comparing the model performance with circular pipe test
data, a preliminary study was undertaken for staticrlly determinate,
reinforced concrete beams with transverse loading and combined trans
verse with axial loading. The purpose of this preliminary study was
to investigate the sensitivity of modeling parameters and to compare
the model predictions with published experimental beam data (8,9) and
conventional ultimate strength theories (4,5). Major findings from
the preliminary study are listed below, additional detail is reported
in Reference (6) .
1. For all the beams studied, including both single and double
reinforcement, the predicted ultimate moment capacity for
transverse loading agreed within 1% to those computed in
accordance with ACI 31877.
2. Predicted loaddeflection curves through ultimate were in
close agreement with experimental data (8) obtained from
two point loading of simply supported, rectangular beams
with approximately 1.7% tension steel reinforcement.
31
3. In the presence of axial thrust loads, the predicted ultimate
moment capacity was in good agreement with experimental data
(9), wherein the ultimate moment capacity initially is in
creased as the axial thrust increased up to the balance
point on the ultimate moment thrust interaction diagram.
Thereafter, the moment capacity steadily decreased to zero
as thrust was increased to ultimate.
4. As expected, the predictions for ultimate thrustmoment
capacity were not influenced by the model input parameters
e , e , and e' which describe the concrete stressstrain
t y c
curve up to compressive strength. Only the strength para
meters for concrete and steel (f ' and f ) influenced ultimate
c y
capacity. However, the loaddeflection path to ultimate is
influenced by e , e , and e 1 and the initial elastic moduli
J t' y* c
values for steel and concrete,
5. The concrete cracking strain parameter e was found to have
a significant effect on the loaddeformation curves for
lightly reinforced beams (typical for culvert crosssections).
As the parameter e decreases over a practical range (0.0001
to 0.0) the effective stiffness decreases resulting in
greater deformations for the same load.
6. The compressive concrete strain parameters, e , and e', also
influence the shape of the loaddeformation curves, but to
a lesser extent than e . As e is decreased over the range
t y
0.0008 to 0.0003 the deformations slightly increase. Con
versely, as e' is decreased over the range 0.0025 to 0.0015
c
deformations decrease.
These preliminary studies demonstrated that the reinforced concrete
model was working properly and provided insights for modeling and
interpreting results for the circular pipes in threeedge bearing dis
cussed next.
32
4.2 EXPERIMENTAL TESTS
The outofground test results used in this study were obtained from
an experimental study by Heger and Saba (15), wherein they tested rein
forced concrete circular pipes under threeedge bearing loadings as shown
in Figure 4.1. Test results included; ultimate strength (load capacity),
0.01 inch cracking l:>ad, deflections, visual observations of crack devel
opment, and stresses; in the reinforcing steel and in the concrete wall.
The pipe test program consisted of 39 pipe specimens with different
wall and diameter dimensions and amounts of reinforcement. For some
pipes, stirrup reinforcement was used to prevent diagonal tension failure.
The unconfined compressive strength of concrete was obtained using
cylinder and core tests, the tensile strength of concrete was obtained
with a split cylinder test, and the ultimate tensile strength, yield
strength and modulus of elasticity for the steel wires were obtained
with tests carried out in accordance with the ASTM Specification A185
56T for Welded Steel Wire Fabric.
From the 39 pipes tested a subset of seven pipes are selected for
this study. The subset represents the complete range of pipe dimensions
and amounts of steel reiinforcement used in the test program. Table 4.1
along with Figure 4.1 identifies the geometry of each selected pipe in
three diameter groups; 48 inch, 72inch and 108 inch pipes (1.22 m, 1.83 m
and 2.74 m) . Each dianeter group has a constant wall thickness with
different amounts of steel reinforcement. Ideally, each group should
consist of low, medium, and high levels of steel reinforcement. However,
the experiment did not include tests with medium levels of reinforcement
for the 43 inch or 108 inch pipe. Thus all groups contain low and high
reinforcement levels, but only the 72inch pipe also has medium reinforce
ment. The first four columns of Table 4.2 shows measured strength properties
of concrete and steel.
33
Load P
!
Oi
//
~1
\\
P/ 2 t fP/ 2
Figure 4.1  Typical Cross Section of Circular Pipe.
Q
I!
>
IT3
Q
Figure 4.2  Finite Element Model of Circular Pise,
34
TABLE 4.1  Geometric Characteristics of the Analyzed
Pipes to Compare Test and CANDE Results
Pipes
Di
ti
Asi
Aso
tbi
tbo
a
(in)
(in)
(in 2 /in)
(in 2 /in)
(in)
(in)
(in)
J
48
5
.01683(L)
. 01233 (L)
1.10
1.09
2
K
48
5
.02708(H)
.01992(H)
1.13
1.11
2
B
72
7
. 03142 (L)
. 02342 (L)
1.14
1.12
3
*
G
72
7
. 05158 (M)
. 03692 (M)
1.18
1.15
3
D
72
7
.07292(H)
.05158(H)
1.21
1.18
3
*
Q
108
9
.05158(L)
. 04000 (L)
1.18
1.16
4.5
p
108
9
.10317(H)
.07383(H)
1.18
1.15
4.5
They have stirrup reinforcement
1 in = 2.54 cm
TABLE 4.2  Material Properties Obtained From
Tests (15) for the Pipes to be Analyzed
Pipes
f
c
ft
*
fsu
fsy
V
c
■k*
fs
(psi)
(psi)
(psi)
(psi)
(psi)
(psi)
(cylinder)
average
average
1
J
4470

81800
79250
4600
80525
2
4730

1
4900
503
79400
77000
5225
78200
K
2
5550
465
1
4900

87300
82000
4400
84650
B 2
4120

3
4640

4
3950

1
4765

88650
85000
4760
86825
G 2
4375

3
5136

' \
6090
—
86100
81500
5820
83800
5550
—
1
5085
507
79100
75500
5810
77300
Q 2
6540
578
1
5175
555
87325
85000
5095
86160
P 2
5015
568
The average from the inner and outer reinforcement
**
The average between the ultimate and yielding stresses
1 psi = 6.895 kPa
36
4.3 ANALYTICAL MODEL AND COMPARISON OF RESULTS
The circular pipe is idealized using the finite element model
shown in Figure 4.2 composed of eleven beam rod elements. For each
of the seven pipes selected there are two or more test results using
the same pipe with the same amount of reinforcement, where some of the
material properties were obtained for each repeated test as shown in
Table 4.2. For analytical predictions, concrete compressive strengths
f from repeated tests are averaged. The value of the steel yield
stress used for analysis is taken as the average between the ultimate
and yielding stresses obtained from the tests. Averaging the ultimate
and yield stress of the reinforcement permits considering both ultimate
load as well as the loaddeflection curve within the limits of perfect
plasticity. The last two columns in Table 4.2 show average strength
values for concrete and steel used for analysis.
Except for the cracking strain parameter e , the remaining material
parameters for steel and concrete are assigned the standard values
(Table 3.1 and 3.2). Since cracking strain is a sensitive parameter
and not well established from the test data, two values are assumed
for analysis; 0.00003 and 0.00008, under the assumption that actual
values will be within this range.
In the following, the analytical predictions (CANDE) are compared
with experimental results for loaddeformation, cracking load, and
ultimate load.
LoadDeformation . Figures 4.3 to 4.16 show predicted and measured
vertical and horizontal deflections versus the applied load for each of
the seven pipes. Each plot shows at least two "repeated" experimental
tests, two predicted curves representing e = 0.00003 and 0.00008, and
the actual mode of failure; flexural or shear. Overall it is observed,
the CANDE predictions generally bracket the experimental curves and
follow the deformation trends quite well. Results are generally in
better agreement when ultimate failure is in flexure rather than shear.
For shear failures, the predicted deflections are generally less than
37
A Load P (Kips)
GO
CANOE (5^.00008)
• CANOE (E t =;00003)
sb. tests
(flexural failure)
0.8
Vertical Oef lection (in)
Figure 4.3  Vertical Load  Vertical Deflection of Pipe J.
A Load P (Kips)
CANOE (£ t =.0000S)
CANOE (8t=.00003)
lab, tests
^— — ^
0.2
(ftexural failure)
0.4 0.S 0.8
Horizontal Deflection (in)
Figure 4.4  Vertical Load  Horizontal Deflection of Pipe J.
38
A Load P(Kips)
02
QA 0.6
CANOE (6f».C0008)
CANOE (E t 00003)
. lab. tests
(diag. tension failure)
08
Vortical Oeflection (in)
Figure 4.5  Vertical Load  Vertical Deflection of Pipe K.
A Load P (Kips)
60
50
40'
30
20
10
CANOE (6 t =.CCC08)
CANOE (S+.00C03 )
lab. tests
(diag. tension failure)
OjS m OS
Horizontal Deflection (in)
Figure 4.6  Vertical Load  Horizontal Deflection of Pipe Z.
39
a Load P (Kips)
0.2
CAN0£(6t = « 00OO8)
CAN0E(8t=.0O0O3)
ab. tests
(diag tension failure)
o,4 as as 10
Vertical Deflection (in)
Figure 4,7  Vertical Load  Vertical Deflection of Pipe B.
A Load P (Kips
 CANOE (5t=.00C03)
 CANOE (6 t =.00003)
ib. tests
(diag. tension failure)
0.8 1.0
Horizontal Deflection (in)
Figure 4.8  Vertical Load  Horizontal Deflection of Pise 3.
40
A Load P (Kips
0.2
n r
0.4
CAN0E(£ t =C'COOS)
CAN0£(£t=OCC0o)
lab. tests
(flexural failure, stirrups:
QJS 0S 10
Vertical (Deflection (in:
Figure 4.9  Vertical Load  Vertical Deflection of Pipe G.
A Load P (Kios)
CANOE (SrCCC0S)
CANOE (8 t = CCCCo)
lab. tests
(flexurai failure, STirrucs)
o.a o.4 o.s oQ io"
Horizontal Ce flection (in]
Figure 4.10  Vertical Lead  Horizontal Deflection of Pipe G.
41
A Load P(Kips)
120
— CANDE (£t=00008)
— CANOE (E t =.00003)
ab. tests
(diag. tension failure)
0.1
0.2 0.3 0.4 0.5
Vertical Oeflection (in)
Figure 4.11  Vertical Load  Vertical Deflection of Pipe D.
A Load P ( Kips)
0.1
0.2
0.3
CANOE (6f =00008)
CA NO E(8"t =.00003)
lab. tests
(diac;. tension failure)
OA
05
Horizontal Oeflection (in)
Figure 4.12  Vertical Load  Horizontal Deflection of Pipe D.
42
A Load P (Kips)
140
120
CANOE (£ t =.OOCCS)
CANDE(£t=00003)
ab. tests
(flexural failure, stirrups)
— >
io
Vertical Deflection (in)
Figure 4.13  Vertical Load  Vertical Deflection of Pipe Q.
^ Load P (Kips)
140
CANOE (£f=CCCCS)
CANOE S^CCCCc)
ab. Tesrs
(flexural failure, stirrucs;
0.S 0.8 1.0
Horizcntal C2fec~.cn [in!
Figure 4.14  Vertical Load  Horizontal Deflection of ?ioe Q.
43
160
140
120
100
80
SO
40
20
A Load P (Kips)
CANO£(E t =.00008)
CANOE (6t=O0003)
lab. tests
0.2
04
(diacj. tension failure)
— >
0.6 0.8
Vertical Oeflection (in)
Figure 4.15  Vertical Load Vertical Deflection, of Pipe P.
<^Load P(Kips)
0.2
0.4
CANOE (£t000C8)
CANOE (Ef=0CC03)
lab. tests
(diag. tension failure)
— i : ' >
0.6 08
Horizontal Oeflection (in)
Figure 4.16  Vertical Load  Horizontal Deflection of Pipe P.
44
measured values. This is to be expected since the CANDE model does not
account for reduced stiffness due to diagonal cracking as shear failure
develops.
For design service loads, say to 2/3 ultimate, the predicted
curves with e = 0,00008 correlate more closely with measured data
than predicted curves with e = 0.00003.
Cracking Load . The cracking load in threeedge bearing is defined
here as the applied load on the test pipe at which a 0.01 inch crack
width occurs extending over a foot in length. The socalled Dload for
0.01 inch cracking (D m ) is the above cracking load per foot of pipe
length divided by the inside pipe diameter. ASTM C7666T describes
five strength classes of reinforced concrete pipe in terms of D .
Table 4.3 shows D values; as measured from the experiments, as
predicted from CANDE, and as specified from ASTM C76 for comparable
steel areas. The CANDE predictions (which employ the GergelyLutz
formula in Chapter 3), are for the case e = 0.00008 which better
represents the experimental data. Comparing CANDE predictions with
test results, good agreement is observed overall. The worst case
occurs for the medium size pipe with heavy reinforcement (pipe D) where
the CANDE prediction is 30% higher than measured. If e is reduced,
the CANDE predictions for D are also reduced.
In view of the random nature of cracking and the inherent approxi
mations in the GergelyLutz formula the CANDE predictions are considered
very reasonable. As a point of interest, the predicted steel stress
was approximately 50 ksi (345,000 kPa) for all pipes when the Gergely
Lutz formula predicted 0.01 inch cracking.
The ASTM D_. values are shown only for reference and indicate
bracketing values for the steel areas used in the experimental tests.
Ultimate Loads . Of the seven pipe types considered in this study,
three failed in flexure and four in shear. Two of the pipes failing
in flexure had stirrup reinforcement to prevent shear failure (diagonal
cracking) at an early load. The modes of failure predicted by CANDE
45
TABLE 4.3  Comparison of Dload at 0.01" Crack
Between the Test, ASTM and CANDE Results
PIPES
DLoad (0.01" Crack)
TEST
TEST
(average)
ASTM
(range)
CANDE
(a)
J
2
1190
1250
1220
1350*
1145
(b) .
K 1
2
1810
1810
1350  2000
1771
(b) x
B 2
3
4
1040
1250
1460
1835
1396
1350*
1400
(c) x
G 2
3
2000
1670
1670
1780
1350  2000
2130
(d)
D i
2
2292
2175
2234
2000  3000
2972
(e)
Q 2
1650
1620
1635
1350*
1216
P l
2
2540
1980
2260
1350  2000
2490
* The asterisk implies the D value is less than the minimum
ASTM rating (1350).
46
agreed with observed failure modes when stirrup reinforcement was taken
into account.
Table 4.4 shows the comparison between test results and CANDE
predictions for the applied load at ultimate flexural failure. CANDE
predictions are in excellent agreement with test data. Predicted ultimate
loads in flexure occur when the slopes of the loaddeflection curves be
come flat, indicating a collapse mechanism has formed.
Table 4.5 shows the load comparison for shear failures. Here,
three CANDE predictions are shown based on three empirical formulas to
estimate ultimate shear stress; (1) ACI formula for straight members
(Chapter 3), (b) Theoretical Modification of Committee 326 for pipes (16),
and (c) MIT Correlation Test formula for pipes (16), When the maximum
shear stress predicted by CANDE reaches the value of these empirical
formulas, shear failure is predicted (assuming no stirrup reinforcement).
Of the three predictions, the standard ACI formula correlates best with
experimental data (except for pipe B).
Summarizing this chapter, we conclude the beam rod element is per
forming very well and is capable of predicting the structural responses
of concrete pipe throughout the entire loading history.
47
Table 4.4.  Flexural Failure of Circular Pipes
PIPES
LOAD **
(kips)
LOAD (kips)
(average)
LOAD (kips)
(CANDE) ~
J l
2
41.6
41.2
41.5
40.8
1
G 2
3
106.0
106.0
106.0
106.0
115.2
k
Q X
2
152.2
152.2
152.2
144.0
* with stirrups
** test load on 4 foot pipe lengths
1 kip = 4.48 kN
Table 4.5.  Shear Failure Loads
PIPES
TEST LOAD (kips)
CANDE LOAD (kips)
each
average
*
(a)
(b)
(c)
1
K
2
59.0
60.2
59.6
56.3
45.6
52.8
1
B 2
3
4
62.4
64.2
64.0
61.5
63.0
73.3
60.0
67.2
D
2
65.2
87.2
76.2
78.5
72.0
91.2
1
P
2
142.8
151.2
147.0
154.2
129.6
165.6
(a) v =
2.0 TV
c
(Standard ACI)
(b) V = 1.6 TV + 64 A ,/D. (Committee 326)
c c si 1
(c) v  1.53 JV + 320 A ./D. (MIT Correlation)
C C SI 1
48
CHAPTER 5
EVALUATION OF REINFORCED CONCRETE MODEL FOR
BOX CULVERTS LOADED IN FOUREDGE BEARING
• Like the previous chapter, this chapter continues to examine the
validity of CANDE's reinforced concrete, beamrod element. Here, we
compare CANDE results with experimental data for box culverts tested in
fouredge bearing. The results to be compared include the load for
0.01 inch cracking and ultimate load. The experimental data did not
include loaddeformation histories, consequently these cannot be com
pared. For reference, the comparisons also include the SGH analytical
predictions (12) based on an elastic analysis discussed in Chapter 2.
5.1 EXPERIMENTAL TEST
The experimental results used for the comparison belong to a test
program (13,14) where outofground reinforced concrete box culverts
with welded wire fabric were loaded up to failure. The loading was
applied as shown in Figure 5,1 using a 4edge bearing testing apparatus.
The material properties of concrete were determined using cylinder tests
and core tests for each kind of box, and the mechanical properties of
the reinforcement were determined by tensile tests (19). Three span
sizes of box were tested, small, medium and large with three levels of
reinforcement in each size, low, medium and high. Thus, nine types of
boxes were tested with two repeated tests per box type.
Tables 5.1 and 5.2 together with Figure 5.1 show the measured
geometries and material strengths for each test box where repeated boxes
are labelled A and B. Note that the core tests for f ' are generally
c
higher than cylinder tests for f and the ultimate steel stress is 10 to
20% higher than initial yield stress.
The test program was performed to verify the SGH analysis/design
method (12) which in turn was used to develop the ASTM standard designs
for reinforced concrete box culverts (21),
49
i
u
CSL
t
P/2 P/2
1 tb2 HVt
Li
ASf
1*3
IK
t \
•AS4
tbi
t fpr As 3
P+W
2 2
Figure 5.1  Typical Cross Section of Concrete Bos Culvert.
u a „
•f — f
f2
HH+^
R+t
• • •
r t 2 3 4 5
6
7t
8*
 Thv+s
9*
*d5 .14
J77777
>Vz
10" t
13 12 it l HV+t £
1, S/? + t/2~
t^ 1
Figure 5.2  Finite Element Model of Concrete Box Culvert.
50
TABLE 5.1
Geometric Characteristics of Test Box Culverts
*•*
BOXES
SPAN
S(in)
RISE
R(in)
HH,HV
(in)
(in)
t3
(in)
tb
(in?
tt>3
(in)
a
:in)
8*48
A
B
96
96
48
48
8
8
8.125
8.250
8.125
8.000
1.376
1.126
1.251
1.251
12
12
8*42
A
B
96
96
48
48
8
8
8.188
8.253
8.251
8.188
1.526
1.278
1.214
1.401
12
12
8*418
A
B
96
96
48
48
8
8
8.250
8.250
8.251
8.126
1.229
1.416
1.354
1.229
12
12
6*410
A
B
72
72
48
48
7
7
7.313
7.313
7.251
7.313
1.053
1.240
1.303
1.240
9
9
6*42
A
B
72
72
48
48
7
7
7.375
7.375
7.438
7.438
1.068
1.443
1.006
1.006
9
9
6*422
A
B
72
72
48
48
7
7
7.313
7.438
7.375
7.375
1.086
1.523
1.148
1.148
9
9
4*44
A
B
48
48
48
48
5
5
5.187
5.188
5.375
5.250
1.353
1.385
1.103
1.353
6
6
4*418
A
B
48
48
48
48
5
5
5.188
5.313
5.188
5.188
1.237
1.112
0.862
1.175
6
6
4*42
A
B
48
48
48
48
5
5
5.563
5.376
5.188
5.250
1.756
1.694
1.381
1.173
6
6
t 9> t_ = thickness of top and bottom slabs respectively
** Boxes are identified according to Reference (14): i.e., span(f t)*rise(ft)
design earth cover (ft) for interstate live load.
1 in = 2.54 cm
51
TABLE 5.2 Reinforcement and Material Properties
of the Box Culverts
BOXES
si
(in2/in)
A s2
(in2/in)
s3
(in 2 /in)
(in)
f c (psi)
cylinder
f c (psi)
cores
*
fy
(ksi)
fsu
(ksi)
8*48
A
B
.02492
.02492
.02492
2
4934
4757
12510
5515
72.3
83.65
8*42
A
B
.04325
.03550
.03550
2
5288
4952
4475
5425
82.0
89.58
8*418
A
B
.04325
.04325
.04325
2
5111
5430
5285
5855
80.5
94.95
6*410
A
B
.01450
.02075
.02075
2
6296
5022
7060
7460
85.9
94.40
6*42
A
B
.03550
.03475
.02675
2
5624
5589
6680
6965
85.8
99.43
6*422
A
B
.02358
.03442
.03442
2
5553
5341
5960
7190
86.2
95.16
4*44
A
B
.01117
.01117
.01117
3
5518
7534
6030
6670
78.5
95.80
4*418
A
B
.01117
.01967
.01967
2
7428
7729
7000
6635
77.3
90.93
4*42
A
B
.01600
.02692
.02692
2
5872
6155
5715
6430
82.0
92.73
*This value is an average of the three reinforcements
S = spacing of longitudinal wires
NOTE: A . steel not used
s4
1 in = 2.54 cm
1 psi = 6.895 kPa
52
5.2 CANDE MODEL
Figure 5.2 shows the finite element model for a typical box culvert
test. Because of symmetry only half the box is modeled with 14 beam rod
elements. The element pattern shown was found to be sufficiently accurate
with regard to element lengths. The reaction support (shown at node 14)
is modeled with a triangular element rather than a boundary condition in
order to avoid imposing a moment constraint (a quirk of CANDE) .
Each element cross section is assigned the concrete thickness,
the steel area, and steel area locations as actually reported from the
experiments (Tables 5.1 and 5.2). Haunches at the box corners are
modeled with two corner elements whose thicknesses are increased by
onehalf the haunch dimensions.
For concrete material properties, f* is taken from the core tests
(Table 5.2, except first box) as this is generally more representative
of each test box, than cylinder tests. The cracking strain is assumed
as e = 0.0001 for all box tests based on observing typical concrete
test results. Other concrete parameters are assigned standard values
(Table 3.1).
Steel "yield" stress for the elasticperfectly plastic model is
taken as the ultimate stress reported in the last column of Table 5.2
for each box. Ultimate steel stress, rather than initial yield stress
is assumed because this better approximates ultimate load capacity.
Other steel parameters are assigned standard values (Table 3.2). Within
each test box the steel stressstrain properties for A ,, A _, and A „
si sZ sj
are assumed identical.
5.3 COMPARISON OF MODELS WITH EXPERIMENTS
In the following comparisons for cracking load and ultimate load
(flexure and shear), the "load" refers to the total applied load P per
foot length of test pipe (see Figure 5.1). Each repeated experimental
test is also repeated analytically with the associated variations in
geometry and material properties.
53
Cracking Load . Table 5.3 shows a comparison between test data and
CANDE predictions for the load producing a 0.01 inch crack. These cracks
occur near the centerline on the inside surfaces of the top or bottom
box slabs. The table specifys "top" or "bottom" indicating which slab
the 0.01 inch crack was first observed, and the CANDE prediction cor
responds to that location. Also shown in Table 5.3, are the SGH pre
dictions for cracking loads to serve as a reference.
Overall it is observed the CANDE predictions are very good and
are statistically better than the SGH predictions as shown at the bottom
of Table 5.3, A graphical comparison of the data is shown in Figure 5.3
from where it is seen that CANDE cracking load predictions are slightly
lower in the average* (conservative) than the test data, but only on the
order of 5 to 10%,
As previously discussed, CANDE predictions are semi empirical and
employ the GergelyLutz crackwidth formula. Although not reported here,
the ACI crackwidth formula (4) was tried with CANDE but not found satis
factory in this study.
Ultimate Loads . In loading the 9 pairs of boxes (18 tests) to
ultimate, 10 tests failed in flexure and 8 tests failed in shear
(diagonal cracking). Two pairs of boxes produced a failure of each
kind. Modes of failure predicted by CANDE agreed with observed failure
modes .
Table 5.4 shows the comparison of the ultimate load for flexural
failure between CANDE prediction, the test results, and the SGH analy
tical results. The values calculated by CANDE are in very good agree
ment with test results, and are slightly better than the SGH analytical
results. Figure 5.4 shows graphically the comparison between CANDE and
the test results for ultimate load at flexural failure. CANDE' s overall
results correlate excellently with the test results, with a + 6% error
range.
Table 5.5 shows the comparison of the ultimate load for shear
failure (diagonal cracking) between CANDE prediction, the test results,
TABLE 5.3 — Comparison of CANDE Results with Test and
SGH Results for 0.01 inch Cracking Load
p .
01 (lb/f
t)
BOXES
TEST
CALCULATED
SGH
Report (14)
CANDE
8*48
A
B
Top
Top
9250
11300
7840
9430
9400
9200
8*42
A
B
Bottom
Top
14000
12300
10950
12360
12300
12300
8*418
A
B
Top
Bottom
13000
13500
15650
13442
14700
14200
6*410
A
B
Bottom
Bottom
9500
9500
5830
6220
7200
7500
6*42
A
B
Bottom
Bottom
14500
10500
10640
10650
10900
10800
6*422
A
B
Bottom
Top
15000
12500
12510
11090
12000
11000
4*44
A
B
Bottom
Top
6700
6000
2740
2700
3600
4200
4*418
A
B
Bottom
Top
7000
8000
7770
7090
7800
7500
4*42
A
B
Bottom
Bottom
7800
8500
6940
8380
7500
8600
p .oi test
Average
1.29
1.10
Standard
Deviation
0.43
0.24
P m calc.
. Ul
Coefficient
of Variation
34%
21%
1 lb/ft = 14.6 N/m
55
16
14
P.01 test 12
( K/f t)
+ 8x4 Boxes
• 6x4 Boxes
x 4x4 Boxes
8 10 12 14 16
P.01 calculated
(K/ft)
P.01 (test)
■ = 1.1063
P.01 (calculated)
Roi (test)
P.01 (calculated)
= 1.0475
* Excluding O
Figure 5.3  Comparison of Test and Calculated 0.01 inch
Crack Load.
56
TABLE 5.4  Comparison of CANDE Ultimate Load with Test and
SGH Ultimate Loads for Flexural Failure
BOXES
uf
(lb/ft)
TEST
CALCULATED
SGH
Report (14)
CANDE
' 8*48
A
B
17860
17230
16050
15780
17700
16800
8*42
A
B
29690
28200
27600
8*418
A
B
6*410
A
B
16100
15000
15380
15390
15600
15600
6*42
A
B
6*422
A
B
4*44
A
B
8980
8440
9800
9011
9600
9600
4*418
A
B
13150
13170
12680
12730
13200
13200
4*42
A
B
19300
18510
18000
P _ test
uf
•
Average
1.03
1,01
Standard
Derivation
0.06
0.06
P calc.
uf
Coefficient
of Variation
6%
. 5.8%
1 lb/ft
14.6 N/m
57
30
+ 8x4 Boxes
•• 6x4 Soxes
x 4x4 Boxes
10 12 14
16 18 20 22 24 26 28 30
Pllf calculated
(K/f t)
^PUf (test)
2PUf (calculated)
1.0129
:P0f (test:
=.9897
'PUf (calculated)
4fr Excluding
Figure 5.4  Comparison of Test and Calculated Ultimata Flexural Load.
58
TABLE 5.5  Comparison of CANDE Ultimate Load with Test
and SGH Ultimate Loads for Shear Failure
Pu dt (lb/ft)
BOXES
TEST
CALCULATED
SGH
Report (14)
CANDE
8*48
A
B
8*42
A
B
22520
21520
22000
8*418
A
B
20890
24490
21590
22860
21600
23000
6*410
A
B
6*42
A
B
19400
25250
23420
23950
22800
24000
6*422
A
B
25'680
21150
21260
23530
21600
23400
4*44
A
B
4*418
A
B
4*42
A
B
14080
12790
13200
Pu,
dt test
Average
1.02
1.01
Standard
Deviation
0.12
0.10
p. ,
dt cal
c.
Coefficient
of Variation
12%
10%
1 lb/ft = 14.6 N/m
59
30
28
26
Pu dt test 24
+» /
(K/ft) 22 .
20
18
16
14
x/
12
h 8x4 Boxes
•■ 6x4 Boxgs
10
x 4x4 Boxes
10 12
8 20 22 24 26 28 30
PlU+ calculated
'dt
(K/ff)
Pbdt ^est)
Pu^t (calculated)
•= 1.0108
Figure 5.5  Comparison of Test and Calculated Ultimate Diagonal
Tension Load.
60
and the SGH analytical results. The values obtained from CANDE are
assuming that the maximum shear stress resisted by the concrete is
2.0 / f f . Once again the values obtained from CANDE are very close to
c
the test results and are a better prediction than the analytical results
of SGH. The test results are compared graphically with CANDE results
in Figure 5.5, from where we can observe that the amount of error from
CANDE is in the range of + 10%, a very good correlation for practical
purposes.
The performance of the beamrod element used in the CANDE program
to model reinforced concrete box culverts has been shown to perform
very well in outofground loading. Subsequent studies will consider
the box culvert buried, subjected to soil loads as well as live loads.
The empirical formulas for crack prediction and shear resistance used
in CANDE will be the same ones used in this chapter, where their per
formance was found satisfactory.
61
CHAPTER 6
DEVELOPMENT OF LEVEL 2 BOX MESH
The reinforced concrete box culvert model is now considered for
its actual function as a conduit buried in soil. Accordingly, both the
soil and the box form the structural system, hereafter called boxsoil
structure. The soil plays a dual role; on the beneficial side it adds
substantial stiffness to the boxsoil structures, on the detrimental
side it transits gravity and applied loads to the box during the in
stallation process.
To determine loads acting on the box requires a complete model of
the boxsoil structure simulating the entire installation process. In
this chapter a general finite element model of the boxsoil structure
is presented with the intent of developing an automated finite element
mesh subroutine suitable for simulating the vast majority of boxsoil
installations encountered in practice. This is called the level 2 box
option of CANDE.
To develop an automated finite element mesh requires some limiting
assumptions and specifications of a variety of parameters describing
the boxsoil system. Overall assumptions are symmetry about the ver
tical centerline and plane strain geometry and loading. Adjustable
system parameters include; box dimensions (span and rise), soil boundary
dimensions (width from centerline, depth below box, and height of cover
above box), soil zones (in situ soil, fill soil, and bedding soil) and
installation type (embankment or trench). These parameters along with
the question of mesh refinement are discussed in the following.
6.1 PARAMETERS TO DEFINE THE MODELS
The depth and width of the entire soil zone are specified in
terms of the particular box dimensions being analyzed. The box culvert
62
is idealized with beam elements located along the middle line of the
walls, so the nominal box span used in our model is equal to the inside
span of the box plus its thickness. Likewise for the nominal rise of
the box. Defining Rl as half the nominal span and R2 as half the
nominal rise, as shown in Figure 6.1, the soil depth below the box is
set at 3R2, and the soil width is set at 4R1 from the box sides. These
soil boundaries are adjudged to be outside the zone of soilstructure
interaction based on previous studies (1).
The height of cover (see Figure 6.1) is an input parameter denoting
the final fill height above the box. However, the height of the mesh
over the top of the box is limited to 3R2 or the specified height of
soil cover, whichever is less. For cover heights greater than 3R2
equivalent loading is used as discussed subsequently.
Other geometry parameters that need to be defined are the trench
depth and trench width as shown in Figure 6.1. If the mesh model is
intended to represent an embankment installation, the trench width is
4R1 so that only fill soil exists on the sides and in situ soil is
leveled with bottom of the box.
The material zones are in situ soil, bedding and fill soil, where
each zone can be assigned the same or different soil mechanical pro
perties. In addition to Rl and R2, the box culvert geometry is defined
with the side, bottom and top slabs thicknesses, and haunch dimensions
as shown in Figure 6.2. The amount of steel reinforcement around the
box is defined by steel areas A , A , A and A along with a common
cover thickness as shown in Figure 6.3.
So far only the parameters of the boxsoil systems and the general
dimensions have been discussed, nothing has been said about the finite
element mesh itself or the sequence of loading. The example shown in
Figure 6.4 will be used to explain the mesh arrangement. All the di
mensions and height of soil cover are shown. A trench configuration
is used with three zones of soil for the system. The values of Rl and
R2 give the overall size of the mesh as previously discussed (see
63
A
CM
cr
00
4 R1
Figure 6.1  Parameters to Define Buried Concrete Box Culvert.
64
Z R2
i
* ' *PTS
Figure 6.2  Parameters to Define the Geometry of the Bos Culvert,
AS4
XL1 kL
XU "R1
Figure. 6.3  Parameters to Define the Reinforcement of the Box Culvert.
65
3R2
3R2
4 R1
Figure 6.4  Example of Box Culvert to Define a Mesh.
66
Figure 6.1). The finite element mesh configuration for the example is
shown in Figure 6.5, where by symmetry only half of the boxsoil system
is modeled. The same figure shows the soil elements and the nodal points
of the mesh. The number of layers of soil elements on top of the box
can decrease if the height cover of soil is less than 3R2, but will
never be less than two rows. The soil elements near the box are smaller,
so that a more refined mesh around the box can provide a better behavior
of the boxsoil model where the stress gradients are known to be highest.
The soil elements are four node, nonconforming quadrilaterals with ex
cellent performance characteristics (1). The coordinates of the nodes
are all related to Rl and R2 and, if desired, can be changed using the
extended level 2 option (see Appendix B).
The loading sequence, called incremental construction (1,2), simu
lates the actual installation process of placing soil layers in a series
of lifts. Figure 6.6 shows the construction increment numbers of element
groups entering sequentially into the system. The initial system (first
construction increment) includes all in situ soil, bedding, and the box
loaded with its own body weight. Subsequent increments, numbers 2
through 9, are gravity loaded layers of fill soil. For specified heights
of soil cover less than 3R2, the mesh over the top of the box is assigned
proportionally less soil layers. If the height of soil cover is greater
than 3R2, the load due to the soil over 3R2 is applied as equivalent
overburden pressure increments. This load sequence is applied after
the ninth soil layer using n9 additional load increments, where n is
the total number of construction increments specified in the input.
The boxsoil mesh described here is generated automatically using
the CANBOX subroutine to generate all the necessary data required to
define the finite element mesh of the system. CANBOX subroutine para
meters, options and mesh size are discussed with more detail in Appendix
B.
6.2 ASSUMPTIONS AND LIMITATIONS
When using this automatic mesh generation, there are some assump
tions involved that should be remembered.
67
MOCfiL MUM3£?.S
137
146
1S3
147
159
150
I SI
152
143 143 ISC iSl
153
152
154
153
ISS
1S4
155
ISS
135
135
137 133
124
1 25
113
114
125
113
139
127
123
140
129
141
142
130
131
115
102 103 104 105
So 57 S3 39
4S
46 47
117
108
113
113
120
107
103
109
gs
95
98
39
81
32
74
67
50
43 43
97
93
90
91
75 75
S4
77
63
r
51
52
70
SO SI
S2
143
132
121
110
99
92
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73
71
S4
53
,44
133
122
111
100
93
36
79
72
S4
34 35
35 37
33 39 40
41
42
43
23
24
25
12
27
23
23
13 14 IS 15
17
30
13
13
12 3 4 3 5
31
20
32
21
10
11
Figure 6,5  Under ormed Grid with Nodal Points.
68
INCREMENT NUneERS
cfc
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7T"
rrr
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"7T
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7T
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Figure 6.6  Load Increment: Pattern for the Layers of Soil.
69
a) To simplify the mesh, we are assuming symmetry about the
vertical axis so only half of the boxsoil system is analyzed. This
assumption implies that only symmetric loading can be applied to the
box when using concentrated loads on top of the mesh. For most of
the cases a symmetric loading arrangement satisfies the loading cond
ditions .
b) Three different zones of soil can be specified within the
soil mesh; fill soil, bedding soil, and in situ soil.
c) When a trench condition is specified, the in situ soil
forming the trench extrados goes all the way up to the top of the
mesh, whereas for the embankment condition, the in situ soil remains
at the level of the bottom slab of the box culvert.
d) For the concrete box culvert, the thickness of the wall is
constant along a particular side but may vary between sides. However,
the concrete cover of the reinforcement is the same for all sides.
Many of the above assumptions can be removed by use of extended
level 2 option (see Appendix B ) which allows selective modification
of the automated mesh discussed above. Virtually all limitations can
be removed by use of level 3 option wherein the user defines his own
mesh (1,2).
70
CHAPTER 7
EVALUATION OF CANDE BOXSOIL SYSTEM
In the previous chapters the reinforced concrete beamrod element
was evaluated with experimental data for outofground structures from
which we concluded that the beamrod element itself performs satisfact
orily. In this chapter we examine the performance of the reinforced
concrete model as a buried box culvert, where the soilbox structure
is modeled with the level 2 box finite element idealization described in
the previous chapter. Using this structural system the loads acting on
the box are not prespecified, but rather are determined from the finite
element solution of the boxsoil system. Thus, the performance of the
box culvert model depends, in part, on the responses of the soil system.
To evaluate the CANDE boxsoil model we first consider a parameter
sensitivity study to assess the influence of soil stiffness and installation
type on the structural behavior of a typical box culvert. Secondly, we
compare the CANDE predictions with full scale field test data (24), pro
viding a direct validation of the boxsoil model.
7.1 SENSITIVITY OF SOIL PARAMETERS
In this section the influence of soil parameters on the structural
performance of a particular box section is examined. Soil parameters
considered include; elastic properties and type of installation (trench
or embankment).
The particular box section (hereafter called standard box) used to
examine sensitivity of the soil parameters was obtained from the ASTM
Standards (21) for box sections under earth dead load conditions. An
intermediate size box with medium reinforcement was chosen to be repre
sentative for this study. Specifically, the standard box has 8 feet (2.4 m)
71
span, 6 feet (1.8 m) rise and 8 inches (20.3 cm) wall thicknesses (8*68)
with a specified design earth cover of 10 feet (3.05 m) . The material
properties for the standard box culvert are:
f
c
and
TC
S„
= 5000 psi
(3A500 kPa)
= 0.0001
= 65000 psi
(448000 kPa)
= 150 pcf
(23.5 kN/m3)
= 4286.8 ksi
(29550 MPa)
= 0.017
= 29000 ksi
(200000 MPa)
= 0.30
1.25 inch
(3.18 cm)
2.0 inch
(5.08 cm)
(unconfined compressive stress of concrete)
(maximum tensile strain of concrete)
(yield stress of reinforcement)
(unit weight of concrete)
(concrete Young's modulus)
(concrete Poisson's ratio)
(steel Young's modulus)
(steel Poisson's ratio)
(concrete cover)
(spacing longitudinal reinforcement)
The characteristics of the 8*68 box cross section are shown in
Table 7.1, where the nomenclature is referred to the typical cross section
shown in Figure 7.1.
For the purposes of this study, elastic soil properties are assumed
in a range covering stiff, medium and soft soils, where their parameters
are Young's modulus and Poisson's ratio. Table 7.2 summarizes the pro
perties of the in situ, bedding and fill soil for the soft, medium and
stiff soil model. The fill soil weight density is assumed 120 pcf (18.8
kN/m*) for all types so that only stiffness is varied. Bedding and
in situ soil zones are not assigned a weight density since they form the
initial configuration.
Installation type . For the first study the influence of installation
type on the standard box is considered for a trench condition versus
72
CM
<r
CM
Figure 7.1  Typical Cross Section and Parameters co Define a
Concrete Box Culvert.
Figure 7.2  Definition of Vertical Deflection o and Horizontal
Deflection 5,
v
a*
73
TABLE 7.1 Characteristics of Standard Box Culvert
Used in Sensitivity Study
Box
(ft*ftin)
ASTM Design
Earth Cover
(ft) .
si
(in2/in)
s2
(in 2 /in)
s3
(in2/in)
A A
s4
(in 2 /in)
XL1
PT
(in)
8*68
10.0
0.01667
0.02417
0.02583
0.01583
0.50
8.0
From Table 3 of ASTM Standards (21).
TABLE 7.2 Properties of the Linear Soil Models
Used in CANDE Solution
Type
*
of Soil
Young ' s Modulus
(psi)
Poission's
Ratio
Insitu (1)
333
0.33
SOFT
Bedding (2)
666
0.33
Fill (3)
333
0.33
Insitu (1)
2000
0.33
MEDIUM
Bedding (2)
4000
0.33
Fill (3)
2000
0.33
Insitu (1)
3333
0.33
STIFF
Bedding (2)
6666
0.33
Fill (3)
3333
0.33
Unit weight is 120 pcf for fill soil
1 ft = 0.3048 m
1 in = 2.54 cm
1 psi = 6.895 kPa
1 pcf = 157.1 N/m"
74
an embankment condition. The trench width beyond the box sides is taken
as 2.0 feet (0.61 m) (narrow trench) and soil properties for both instal
lation types are assigned medium stiffness values (see Table 7.2). In
both cases, the box was loaded up to 28 feet (8.53 m) of soil cover above
the box. Figure 7.2 shows the definition of vertical and horizontal
relative displacements used in subsequent discussions.
Figure 7.3 shows the fill height versus vertical displacement history
of the box, from where it's observed that the embankment configuration
produces slightly greater vertical deflections in the box. Figure 7.4
shows the bending moment diagrams and shear force diagrams in the box
at 28 feet of soil cover for both installation types. The embankment
condition gives greater bending moments and shear forces acting in the
box, which conforms to the greater deflections previously observed.
From this comparison it was concluded that the embankment condition
produces slightly greater loading conditions on the box so that in all
subsequent studies presented herein only the embankment condition will
be considered.
Soil Stiffness . The effect of elastic soil stiffness is investigated
using the same standard box (8*68) with an embankment soil configuration,
where the properties of the soil are varied to idealize a soft, medium
and stiff soil as defined in Table 7.2. With these values a range of
variation is covered so the effect of each can be observed. Figure 7.5
shows the load versus vertical deflection history of the box for the
three classes of soil stiffness, where the box is loaded up to failure
for each case. The failures for the box culverts are defined by exceeding
ultimate shear capacity (V ) or by the formation of plastic hinge mechanisms
from excessive moments and thrust (M ) . In this study shear failure
occurs before plastic hinging in the standard box (8*68) for all the
three types of soil. Failure occurs first for the soft condition, whereas
for the medium and stiff soil conditions failure occurs at a greater
height of soil cover. Note that when ultimate bending failure (M ) occurs,
the deflections do not show a flat slope (increase without bounds) as
75
.Height of So
i (ft)
25
20
15
to
5
— Troncn
— embankment
8x68 Box
medium reinfor.
standard soif
0.1
02
03
0.4
**fv
Vertical Oisplacemen
(in)
Figure 7.3  Height of Soil Over the Top of the Box  Vertical
Deflection for Trench and Embankment Situations.
76
(a) Bending Moment diagram
for trench ■ 2.0 ft.
(b) Bending Moment diagram
for embankment
Cc) Shear Force diagram
for trench 3 2.0 ft.
(d) Shear Force diagram
for embankment
Figure 7.4  Bending Moment Diagrams and Shear Fo'rce Diagrams of
8x6 Box Culvert with 28 ft of Soil Cover for Trench
* 2.0 ft and. Embankment Conditions.
77
Mu
/
A Heiqht of soil /
* (ft) /***,
Vu /
25
20
Mu
Vu
Stiff SOil
medium soil
soft soil
8x68 Sox
medium reinforce:
L—
0.2
0.4
05
0.8
■»f
v
Vertical Displacement
(in)
Figure 7.5  Height of Soil Over Che Top of che Box  Vertical
Deflection for Three Kinds of Soil.
78
they did in previous studies with outofground culverts with applied
loads. This is because the soil stiffness is now controlling deflections
at ultimate. Accordingly, the slope of the deflection curves at ultimate
is in near proportion to soil stiffness. From Figure 7.5 it is evident
that the soft soil condition is restricted to smaller cover heights to
reach ultimate than the stiff er soils. Thus, the medium and stiff soil
conditions are more favorable for the box behavior.
Figure 7.6 shows the bending moment diagrams for the box at bending
failure (presuming stirrups) for the three types of soil conditions. It
is observed that even when the failure occurs at different heights of
soil cover, the maximum moments are similar as would be expected for
ultimate moments. From this comparison it is evident that the type of
soil is an important factor for analysis and design of a box culvert.
7.2 COMPARISON WITH TEST DATA
To validate the boxsoil model, CANDE results are compared with test
data from a full scale field installation. Test data on buried box
culverts is very limited. However, recent research reports from the
Department of Transportation, Lexington, Kentucky (23,24,25) have supplied
some test data. From these reports, data was obtained for a box culvert
in Clark County, Kentucky, designed as an embankment with a yielding
foundation within a bedrock formation (24). Instrumentation on the box
included normal pressure gages and a few strain gages on reinforcement
steel which were reported not to function properly, thus only normal
pressure comparisons are used for this study. The box is identified
as Station 123+95 in the report (24) and its cross section as modeled
by CANDE is shown in Figure 7.7 along with reinforcement areas.
The level 2 box embankment condition is used for the model with
three zones of soil that are assumed linear with the properties shown
in Table 7.3. The in situ soil is bedrock so a large value is assumed
for its modulus of elasticity. Table 7.4 shows the material properties
used in CANDE to model the box culvert. The concrete and steel strengths
used in CANDE were obtained from data presented in a report (25) and the
79
(a) Stiff soil and 31 ft of soil cover
A
189*7
CD
y
/
CO
^^
/
iO — '
I
M
dbin)
CO
\
X
21400
(b) Medium soil and 28 ft of soil cover
(c) Soft soil and 22 ft of soil cover
Figure 7.6  3ending Momenc Diagram of 8x6 Box Culvert at 3ending
Failure Usinc Three Different Kinds of Soil.
80
z
77777777777777777T77/
(solid rock)//// f *
y////////M»
Tfgure7.7 Cross Section of Buried Test Box Culvert
(Station 123+95).
81
TABLE 7.3 
Linear Soil Properties for Test
Box Culverts
SOIL
Young's Modulus
(psi)
Poisson's Ratio
INSITU
(rock)
100000
0.25
BEDDING
4000
0.25
FILL
2000
0.25
Note: soil weight density of 138 pcf for fill soil
TABLE 7.4  Box Culvert Properties
BOX
f
c
(psi)
Y
c
(pcf)
f
y
(psi)
e
t
Station
123+95
4500
150
60O00
.0001
1 psi  6.895 kPa
1 pcf = 157.1 N/nf
82
other properties are assigned standard values (Tables 3.1 and 3.2).
The box is loaded up to 77 feet (23.5 m) of soil cover using small
load increments thereby obtaining a history of the boxsoil system
performance. However, the only information that can be used for com
parison is the pressure distribution on the box which was experimentally
measured at two heights of fill soil, 21.6 feet (6.58 m) and 77 feet (23.5 m) .
To measure the pressure around the box, eight Carlson earth pressure
cells were installed, two on each side of the box. Figure 7.8 shows the
CANDE pressure distribution around the box at the two fill heights of
soil, along with the measured test data. CANDE predictions of the pres
sure is very close to the measured value for the top and bottom slabs.
The measured pressure on the sidewalls is different for the right wall
and the left wall, and CANDE prediction is closer to the values measured
for the right wall.
From CANDE, some interesting observations are, when loaded up to
the maximum 77 feet (23.5 m) of soil cover, the bottom corner steel
started to yield for the last load increment. Also, the 0.01 inch
(0*0254 cm) crack first developed with 60 feet (18.3 m) of soil cover.
These observations suggest an economical design was achieved with no
conservatism.
After this last study it can be said that the reinforced concrete
beam element to model the box culverts and the boxsoil system appear,
to give reasonably good predictions of the behavior and performance of
buried box culverts.
83
fill heiqht 21.6' 1 ^ AKin r «~ ,
J > CANOE normal pressure
fill height 77' J
instant value (H= 21.6')
, > test normal pressure
lapsed time value (H=77) J
SCALE:
50 psi
Figure 7.8  Comparison of Test Data with CANDE Prediction
of Pressure Over the Box Culvert.
84
CHAPTER 8
EVALUATION OF ASTM C789 DESIGN TABLES WITH CANDE
In the previous chapter, CANDE' s reinforced concrete beamrod
element has been developed and compared with experimental data for both
in ground and outofground culverts. Overall, very good correlation
was observed for all aspects of structural performance, including;
loaddeformation curves, cracking loads, ultimate loads, and soil pres
sures, thereby lending a measure of confidence and validity to the
CANDE model.
In this chapter the objective is to crossevaluate CANDE with ASTM
C789 design tables for buried box culverts (21). As discussed in Chapter
2, the ASTM design tables are based on an elastic method of standard
analysis together with the ultimate method of reinforced concrete design
(12) . However, the magnitude and distribution of loads acting on the
box are assumed, as opposed to determining loads with soilstructure
interaction models like CANDE. In Chapter 5 it was shown that CANDE'
predictions for outofground box culverts loaded in bearing correlated
very closely with the analytical predictions subsequently used to develop
the ASTM design tables (14). Thus for buried boxes, it may be presumed
that comparisons between CANDE and ASTM design tables will be influenced
primarily by the modeling of soil and soilstructure interaction as op
posed to the modeling of the box.
The comparisons reported herein are divided into two main sections;
(1) dead loading due to soil weight only (ASTM C789, Table 3), and (2)
dead loading due to soil weight plus HS20 live loading conditions (ASTM
C789, Table 1).
8.1 BOX SECTION STUDIES FOR DEAD LOAD
Table 3 of ASTM C789 lists the design earth cover (allowable fill
height) for each standard box size as a function of the steel reinforce
85
ment areas A ., ^c?* ^cV anc * ^qa' ^ or this study, a subset of these
standard boxes were selected covering the typical range of box spans,
rise/span ratios and amounts of steel reinforcement. These subsets
are shown in Tables 8.1a and 8.1b.
Table 8.1a represents the typical range of box spans; large (10 foot
span) , intermediate (8 foot span) and small (4 foot span) where the
span/rise ratio is an intermediate range 1.3 to 1.7. For each box, three
levels of steel area (low, medium, and high) are listed and correspond
to increased levels of design earth cover. In a similar manner, Table
8.1b identifies three standard boxes with span/rise ratios ranging from
1.0 to 2.0 and a common box span of 8 feet (2.44 m) . Taken together,
Tables 8.1a and 8.1b cover the typical range of the standard ASTM box
designs. Note that the intermediate box 8*68 (span*rise inches wall
thickness) is common to both tables. Thus, there is a total of 5 different
box sizes with three levels of reinforcement, providing 15 different box
sections for comparative analysis.
Comparison Objectives and CANDE Model . For each of the ASTM box sections
defined above, CANDE predictions are compared with ASTM assumptions for
(a) soil load distribution on box at design earth cover, and (b) soil
load distribution on box at failure cover heights. In addition, the
consistency of ASTM designs are evaluated with CANDE with regard to 0.01
inch cracking load and failure load.
In order to make these comparisons, the parameters of the CANDE
model are defined as consistantly as possible with ASTM assumptions.
The concrete properties assumed for each box are:
f£ = 5000 psi (345000 kPa) compressive strength
e t = 0.0001 in/in cracking tensile strain
Y c = 150 lbs/ft 3 (23.5 kN/m 3 ) weight density
The remaining concrete parameters are taken as the standard values
in Table 3.1.
86
TABLE 8.1 Reinforcement of Concrete Box Culverts Under Earth
Dead Load Conditions (ASTM Table 3) Used for Comparison
(a) Span/Rise Approximately 1.5, Span = large, intermediate and small
BOX
ASTM Design
Earth Cover
(ft)
Reinf .
si
(in 2 /in)
s2
(in 2 /in)
A Q
s3
(in 2 /in)
A /
s4
(in2/in)
6
Low
.02000
.02000
.02000
.02000
10*610
10
Medium
.02333
.02833
.03000
.02000
14
High
.03250
.03833
.04000
.02000
6
Low
.01583
.01583
.01667
.01583
8*68
10
Medium
.01667
.02417
.02583
.01583
14
High
.02333
.03333
.03500
.01583
10
Low
.01000
.01000
.01083
.01000
4*35
14
Medium
. 01000
.01417
.01000
. 01000
18
High
.01167
.01833
.01833
.01000
(b) Intermediate Span, Span/Rise from 1.0 to 2.0
BOX
ASTM Design
Earth Cover
(ft)
Reinf.
A i
si
(in 2 /in)
A
s2
(in2/in)
A O
s3
(in 2 /in)
A /
s4
(in 2 /in)
8*48
6
10
14
Low
Medium
High
.01583
.02000
.02833
.01583
.02166
.02917
.01583
.02250
.03000
.01583
.01583
.01583
8*68
6
10
14
Low
: Medium
High
.01583
.01667
.02333
.01583
.02417
.03333
.01667
.02583
.03500
.01583
.01583
.01583
8*88
5
8
12
Low
Medium
High
.01583
.01583
.01750
.01583
.02167
.03083
.01667
.02333
.03333
.01583
.01583
.01583
1 ft = 0.3048 m
1 in = 2.54 cm
07
and
Assumed steel properties are:
f y = 65000 psi (448000 kPa) yield stress
E s = 29000 ksi (200000 MPa) Young's modulus
T c = 1.25 in (3.18 cm) concrete cover to steel center
S^ = 2.00 in (5.08 cm) longitudinal spacing for crack prediction
An example of the CANDE input parameters for 8x6 box is given in
Appendix C.
Since the ASTM approach does not consider soil stiffness, the CANDE
solutions use two soil conditions, soft and stiff, for the analysis of each
box, thereby bracketing the practical range of soil stiffness. Soil moduli
values for soft and stiff conditions are given in previous chapter in Table
3
7.2. For both conditions, soil density is taken as 120 pcf (18.8 kN/m ).
All CANDE solutions are obtained using the new level 2 box generation
scheme for an embankment installation. Nine construction increments of
soil are used tobring the soil height up to the ASTM design cover height
to facilitate the comparison of loading distributions assumed by ASTM with
those predicted by CANDE. Thereafter, additional soil layers are added
until flexural failure is observed. During this loading sequence, the
cover height causing initial 0.01 inch cracking is determined along with
the cover height causing shear failure, providing shear failure occurs
before flexural failure.
Load Distribution Comparisons at Design Cover Height . The ASTM assumed
load pattern due' to soil pressure and box weight are shown in Figure 8.1.
Vertical soil pressures are assumed uniform and proportional to cover
height. Lateral pressures are assumed to vary linearly, dependent on
the coefficient of lateral earth pressure generally assumed to be 0.5.
No shear traction on the box sides is assumed in the ASTM pattern.
Figure 8.2 illustrates, the nature of a typical load distribution
predicted by CANDE resulting from soil loading and box weight. Vertical
soil pressures are not uniform, lower in the middle where bending defor
mation is greatest. Lateral pressure along the box increases with depth
88
Fs
uj c
{ :l t
<Jc
C"s
t ■> ;'
Z y ; t, :, i. j, y I i. Z —
f—s
r  r
m
2 ,i 4 ' + ** T 7 '<
**
H
Ghs
R
1 G"wn
(7s = X* s x H
Oc=W/(S+2t)
aic = !fcx t
(Tms = 0.5(Cs)
(Thu^Ghs* 0.5 3s(R+t)
W = Total Weight of
Box
S =• Span of Box
R = Rise of Box
H = Height of Soil
from. Top of Box
Oc = Unit Weight of
Concrete
#S = IJnit Wei Sat of
Soil
+ s Thickness of Box
0< = Coefficient for
Lateral Pressure
Figure 8.1 Loading of a Box Culvert Due to Soil According
to ASTM Norms.
UJvy=
OJH UJ w =
Sv=
Normal Load in the
Vertical Direction
Normal Load in the
Horizontal Direction
Shear Load Over the
Walls
Figure 8.2 Typical Load Pattern Due co Soil Load Obtained from CANDE.
89
but not linearly. This applies to both sides of the box but is only
illustrated on the right side in Figure 8.2. In addition to the normal
pressures, significant shear traction develops over the side walls
acting mostly downward. This is illustrated on the left side of the box.
Shear traction on the top and bottom slabs is also present but is not
significant and is not shown. Shear traction on the side walls can
amount to 50% of the net downward force which must be equilibrated by
the pressure along the bottom slab. This is an effect not considered
in the ASTM load pattern and should be kept in mind in the subsequent
comparisons.
In order to compare ASTM and CANDE load distributions at design
earth cover, normalized plots are constructed by dividing the CANDE
predictions by the ASTM assumption at each point around the box. This
is clarified in Figure 8.3 where the dashed lines represent normalized
values of unity, and the solid lines represent the ratio of CANDE pre
diction to ASTM assumption. Shear traction is arbitrarily normalized
by dividing the CANDE prediction for shear traction by the ASTM assump
tion for normal pressure on the top slab. Due to symmetry, both sides
on the box experience identical loading distributions. Normalized
plots for lateral soil pressure are shown on the right side of the box,
while normalized plots for shear traction are shown on the left.
With the above understanding, Figures 8.4 through 8.8 show the
normalized load distributions for each box in Table 8.1a,b. Each
figure shows six normalized plots per box representing the three levels
of reinforcement and the two soil conditions.
In general, CANDE predictions for the normal pressure on the top
and bottom slabs are not uniform, increasing from the center of the
slab to the corner of the box. Normal pressure at the center of the
top slab are very close to ASTM assumption and increases to a range of
20% to 30% greater than the ASTM assumption near the corner, depending
on the soil conditions and level of reinforcement. CANDE predictions for
the normal pressure on the bottom slab is significantly higher for soft
soil than stiff soil. This is because soft soil generates greater shear
90
^
Olw =
V= ^v/ffs+UJc)
Figure 8.3  Explanation of the Plots for Normalized Normal
Pressure and Shear Forces Acting on the 3ox
(CANDE Prediction/ ASTM Assumption) .
91
(a) Low reinforced and
stiff soil
UJ V
UJ V
(b) Low reinforced and
soft soil
(c) Medium reinforced and
stiff soil
UJ V
(d) Medium reinforced and
soft soil
UJ V
Hs 14.0 f"
(e) High reinforced and
stiff soil
vu,
UJy
(f) High reinforced and
soft soil
Figure 8.4  Normalized Plots for Normal Pressure and Shear Acting
on a 10x6 Box iZ ASTM Height of Soil
92
(a) Low reinforced and
stiff soil
(c) Medium reinforced and
stiff soil
(e) High reinforced and
stiff soil
Ulu
I UJ,
(b) Low reinforced and
soft soi
PH
(d) Medium reinforced and
soft soil
(f) High reinforced and
soft soil
Figure 8.5  Normalized Plots for Normal Pressure and Shear Acting
ou a 8x6 Box ac ASTM Height of Soil.
93
UJH
(a) Low reinforced and
stiff soil
UJ V
(b) Low reinforced and
soft soil
UJv
UJ V
(c) Medium reinforced and
stiff soil
(d) Medium reinforced and
soft soil
UJ V
(f) High reinforced and
soft soil
(e) High reinforced and
stiff soil
Figure 8.6  Normalized Plots for Normal Pressure and Shear Acting
on a 4x3 Box at ASTM Height of Soil.
94
(a) Low reinforced and
stiff soil
UJ H
I
i
H=6.0 ft
Ivjl
\
{
1
"v "**
(d) Lov reinforced and
soft soil
UJ V
(c) Medium reinforced and
stiff soil
UJh
(d) Medium reinforced and
soft soil
UJ V
UJy_
(e) High reinforced and
stiff soil
ujh
oj v
(f) High reinforced and
soft soil
Figure 8.7  Normalized Plots for Normal Pressure and Shear Acting
on a 8x4 Box at ASTM Height of Soil.
 UJh
95
(a) Low reinforced and
stiff soil
(b) Low reinforced and
soft soil
uj v
UJv
(c) Medium reinforced and
stiff soil
(d) Medium reinforced and
soft soil
UJ V
UJv
(e) High reinforced and
stiff soil
UJv
(f) High reinforced and
soft soil
Figure 8.8  Normalized Plocs for Normal Pressure and Shear Acting
on a 8x8 Box at ASTM Height of Soil.
96
forces over the side walls producing a greater downward force. For the
stiff soil condition CANDE predictions are similar to ASTM assumption at
the center of the bottom slab and increases to about 60% to 70% greater
than the ASTM assumption near the corner. For the soft soil condition
CANDE predictions are about 20% to 40% greater than the ASTM assumption
at the center of the bottom slab and increases to about 70% to 100%
greater near the corners. The lateral pressure from CANDE predictions
are not linear like the ASTM assumption, but in general the magnitudes
are close.
Load Distribution Comparisons at Failure Cover Heights . To further this
study, each box was loaded beyond the ASTM design earth cover to failure
for both soil conditions. Figures 8.9 to 8.13 show loaddeflection histories
of all the boxes analyzed to failure. For each box shown, the type of
failure that first occured is indicated at the height of soil cover where
failure occured. As expected, the boxes buried under soft soil conditions
exhibit greater deflections during their load history and fail prior to
the identical box analyzed in stiff soil conditions.
For fill heights at failure, normalized loaddistribution plots
(CANDE prediction divided by the ASTM assumption) are constructed in
the same manner as previously described and are shown in Figures 8.14 to
8.18. Note, the magnitude of the ASTM load distributions are linearly
related to cover height but retain the same shape for all fill heights.
Load distributions from CANDE change both in magnitude and in shape
during loading as a consequence of soilstructure interaction and
changing stiffness of the box.
The normalized plots at failure show the same general trends as the
normalized plots at design earth cover. Now, however, the normal pressure
distributions on the top and bottom slab tend to increase more rapidly,
beginning with relatively smaller magnitudes at the slab centers and in
creasing to relatively higher magnitudes at the slab corners. This is
attributed to the reduction of slab bending stiffness as failure develops,
i.e., a greater portion of the soil load is shifted to the stiff er corners
where the side walls serve as thrust columns.
97
.0041
.008
.012
.016
.020
.024
&
soft SO i
'/r
o
+
high reinf
medium reinf
low reinf
bonding failure
shear failure
(a) Vertical deflection/rise vs. fill height /rise.
.002
.0041
.006
.0081
Sh
1 2 3
1 i ~
5 6 H /r
iff soil
SOft SOil
'/s
(b) Horizontal deflection/ span vs. fill height/rise.
Figure 8.9 
Height of Soil  Deflection for 10x6 Box with
Soil Loads.
98
3 4 5 H/ R
.0021
,004
.006
.008
.010
.012
H
Stiff SCi
SOft SOi
high re inf.
medium reinf
low reinf
o bending failure
. + shear failure
(a) Vertical deflection /rise vs. fill height /rise.
5 HA
.002
.oo4^
.006
Ws
stiff so
soft so
(b) Horizontal deflection/ span vs. fill height/rise.
Figure 8.10  Height of Soil  Deflection for 3x6 3ox with
Soil Loads.
99
8 10 H /R
SOft SOil
S^^stif f SOii
\ \
\ \
\ \
high reinf
I
medium reinf
1
low reinf
i
bonding failure
+► shoar failure
.002
.004
.00 G
.008
.010
.012
Sv /R
(a) Vertical deflection/ rise vs. fill height/rise
.001
.002
£031
ft
4 /s
Stiff SOil
10 H/ R
soft so
(b) Horizontal deflection/ span vs.. fill height /rise.
Figure 8.11  Height of Soil  Deflection for 4x3 Box with
Soil Loads.
100
10 h /r
.030
high rsinf \
m odium rsinf j
low roinf? \
o bonding failursj
+ shoar failure
9
(a) Vertical deflection/rise vs. fill height/rise.
.002
,004
.006
r.
%
Stiff SOi
10 H/
R
soft so
(h) Horizontal deflection/ spaa vs.: fill height/rise.
Figure 8.12  Height of Soil  Deflection for 8x4 Box with
Soil Loads.
101
2 3
h /p
.002
.004
.006
.008
.010
.012
■ , ~^ g ^' ' L "
SOft SOil^ \\ \
V
\
\
"\
I
i 1
i ' ii
^rStif f SOil
high reinf
medium reinf.
low reinf.
o bonding failure
+• shear failure
(a) Vertical deflection/rise vs. fill height /rise,
1 2 3 4 5 H / R
soft soi
stiff soil
.001
.002
.003
fys
(h) Horizontal deflection/span vs. fill height /rise.
Figure 8.13  Height of Soil  Deflection for 8x8" Box with
Soil Loads.
102
UJv
UJ V
(a) Low reinforced and
stiff soil
ULly
(b) Low reinforced and
soft soil
(c) Medium reinforced and
stiff soil
UJv/
(d) Medium reinforced and
soft soil
(e) High reinforced and
stiff soil
(f ) High reinforced and
soft soil
Figure 8.14  Normalized Plots for Normal Pressure and Shear Acting
on a 10x6 Box at Failure Load.
103
(a) Low reinforced and
stiff soil
(c) Medium reinforced and
stiff soil
(b) Low reinforced and
soft soil
UJy
(d) Medium reinforced and
soft soil
Note: The 8x6 box with high reinforcement has plots similar to the medium
reinforced, where the failure is due to shear.
Figure 8.15  Normalized Plots for Normal Pressure and Shear Acting
on a 8x6 Box at Failure Load.
104
(a) Low reinforced and
stiff soil
(c) Medium reinforced and
stiff soil
(b) Low reinforced and
soft soil
(d) Medium reinforced and
soft soil
Note: The 4x3 box with high reinforcement has plots similar Co the
medium reinforced, where the failure is due to shear.
Figure 8.16  Normalized Plots for Normal Pressure and Shear Acting
on a 4x3 Box at Failure Load.
105
UJ V
uu v
rr
i
\
\
H=22.0 ft .
1
1
jS v
1
1
1
3/^
(a) Low reinforced and
stlf
f soil
UJh
uj h
(b) Low reinforced and
soft soil
(c) Medium reinforced and
stiff soil
UJ V
(d) Medium reinforced, and
soft soil
Noter The 8x8 box with high reinforcement has plots similar to the
medium reinforced, where the failure is due to shear.
Figure 8.17 Normalized Plots for Normal Pressure and Shear Acting
on a 8x8 Box at Failure Load.
106
(a) Low reinforced and
stiff soil
(b) Low reinforced and
soft soil
(c) Medium reinforced and
stiff soil
(d) Most reinforced and
soft soil
(e) High reinforced and
stiff soil
Ol v
(f ) High reinforced and
soft soil
Figure 8.18  Normalized Plots for Normal Pressure and Shear Acting
on a 8x4 Box at Failure Load.
107
Lateral pressure on the side wall tend to be greater than ASTM
predictions, particularly in the center region where outward deflections
mobilize passive soil resistance. Side wall shear traction is maximum
at the top walls acting in the downward direction. Near the bottom,
shear traction reverses sign, but the net effect is a significant down
ward force, an effect not considered in the ASTM assumed load pattern.
Cracking and Failure Loads Comparisons . The CANDE prediction for the
fill height producing 0.01 inch cracking is shown in Tables 8.2 and 8.3
(second column from end) corresponding to the boxes defined in Tables 8.1a
and 8.1b for both soft and stiff soil conditions. For each box size it
is observed that the cracking load (fill height) increases with the
level of reinforcement. Also for identical box cross sections, the
cracking load increases with soil stiffness.
To check if the ASTM design earth covers are conservative compared
to predicted fill heights at which 0.01 inch cracking occurs, a fill
height ratio (CANDE prediction/ASTM design cover) is shown in the last
columns. This ratio should be more than 1.0 for conservative designs.
For stiff soils, the cracking load ratio varies from 1.06 to 2.02 im
plying the ASTM design covers are conservative. For soft soils, the
ratio varies from 0.75 to 1.44 implying some designs may not be conser
vative. Boxes with low reinforcement tend to be more conservative than
with high reinforcement. As a general conclusion, the ASTM boxes are
moderately conservative with respect to 0.01 inch cracking at design
earth cover providing good quality soil is used.
Also shown in Tables 8.2 and 8.3 are the CANDE predictions for
fill height at failure as controlled by flexure or shear. In most
cases shear failure occurs prior to flexure failure except for some
lightly reinforced boxes. Identical boxes fail at lower fill heights
in soft soil than in stiff soil. Both shear and flexure failure heights
are reduced in soft soils, but flexure failure heights are reduced by
a greater precentage.
A "failure load ratio" is defined here by dividing the predicted
failure height (as controlled by shear or flexure) by the ASTM design
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earth cover and is tabulated in the center column of the tables. Pre
sumably, the ASTM designs are based on a load factor of 1.5 times the
design earth load. Thus, the failure load ratio defined above should be
at least 1.5 to achieve the intended ultimate capacity. For stiff soils,
the ratio varies from 1.5 to 4.4, whereas for soft soils, the ratio
varies from 1.3 to 3.6. For a given soil condition and box size, the
ratios are higher for low reinforcement than for high reinforcement.
Overall, it is concluded that the ASTM box designs are conservative with
respect to the 1.5 load factor criterion when good quality soil is used,
but less so for the high reinforcement than low reinforcement. In other
words, the ASTM design earth cover specified for a box with low reinforce
ment is more conservative than the specified earth cover for the identical
box with high reinforcement.
8.2 BOX SECTION STUDIES WITH LIVE LOADS
In this section the effect of live loads on shallowly buried boxes
are investigated and compared with ASTM C789 design tables.
The ASTM Specifications consider two types of live load in their
box culvert design tables, HS20 truck loads (ASTM Table 1) and inter
state truck loads (ASTM Table 2). Due to the small difference between
these design tables, only the HS20 live loads are considered in this
study. Figure 8.19 shows the HS20 truck axle loads along with an
"equivalent" transverse strip load P = 222 lb/in (389 N/cm) used as a
reference plane strain loading in the CANDE analysis. The strip load P
represents the static weight of the middle axle tire loads distributed
by the axle length as shown in the Figure.
The boxsoil system analyzed in CANDE is an embankment installation
using the Level 2 box automatic mesh generation along with the extended
level 2 option for defining the live loads. As before, two types of soil,
stiff and soft (see Table 7.2), are used for each box culvert. Figure 8.20
shows a typical box culvert cross section, where the live load P repre
senting the HS20 truck's middle axle is applied over tne center of the
box culvert. The other axles are well away from the box culverts con
Ill
B
8000 lb 320001b 320001b
f^4
/ f r\*
14=30
f^f
HS20 Live Load
f^4
\,
5 ISOOOIb 160001b
I 1
kiu hud jSl*^
p= 32000 lb =222 i b/
144 in
«n
Figure 8.19  HS20 Truck Live Load and Equivalent Plane Strain
Strip Load.
P=222 1 b/ in
(/ns/t
4
Figure a. 20  Typical Box Culvert Cross Section Used for Live Load
Comparison Study.
112
sidered herein and have negligible influence on the box deformations.
To study the effect of live load at shallow soil cover, a set of
five boxes representing a range of sizes and span/rise ratios were
selected from ASTM Table 1. Each of these boxes has as a minimum
allowable fill height of 2.0 feet (0.61 m) to disperse the concentrated
live load. Table 8.4 lists these boxes and the ASTM steel reinforcement
areas specified for the common minimum soil cover. For the CANDE analysis
each box was incrementally loaded with soil layers up to 2.0 feet
cover height. Next, the live load P was applied and then incrementally
increased to a value P* at which a 0.01 inch crack occurred in the box
culvert. This was repeated for each of the five boxes using stiff and
soft soil conditions.
By forming the load ratio P*/P, the ASTM designs can be evaluated
on the basis of 0.01 inch cracking design criterion. Table 8.5 shows
the results of this study. Note that the load ratios are high but
fairly uniform, ranging from 4.0 to 4.7 including both soil conditions.
Thus, it is concluded that ASTM designs at 2.0 ft cover heights tend to
be overly conservative even if impact loads are added to the HS20
loading (ASTM recommends impact loads up to 20%) .
Studies similar to the above indicated that the influence of live
loads is negligible compared to dead loads for fill heights greater
than eight feet.
113
Table 8.4 Reinforcement of Box Culverts with Minimum
Soil Cover and HS20 Live Load (ASTM Table 1)
BOX
ASTM
H soil
(ft)
. A si
(in 2 /in)
A S2
(in 2 /in)
A S3
(in 2 /in)
A S4
(in2/in)
4*35
2.0
.01750
.02250
.02000
.01000
8*68
2.0
.02583
.03833
.02917
.01583
10*610
2.0
.02917
.03833
.02833
.02000
8*88
2.0
.02167
.04250
.03333
.01583
8*48
2.0
.Q3Q83
.03333
.02417
.01583
Table 8.5 Live Load Performance Factor for Box Culverts
with Minimal Soil Cover and HS20 Live Load
BOX
ASTM
H soil
(ft)
HS20
P
(lb/ in)
Type of
Soil
*
P
(lb/in)
P*/P
4*35
2.0
222
STIFF
SOFT
1046
942
4.71
4.24
'8*68
2.0
222
STIFF
SOFT
1006
942
4.51
4.24
10*610
2.0
222
STIFF
SOFT
1030
930
4.64
4.19
8*48
2.0
222
STIFF
SOFT
1022
892
4.60
4.02
8*88
2.0
222
STIFF
SOFT
926
880
4.17
3.96
*Result obtained from CANDE
1 ft = 0.3048 m
1 lb/in = 1.22 N/m
114
CHAPTER 9
SOIL MODELS
Soil models originally incorporated into the CANDE1976 program
included: (a) linear elastic (isotropic or orthotropic) ; (b) incremental
elastic, wherein elastic moduli are dependent on current fill height
(overburden dependent); and (c) variable modulus model using a modified
version of the Hardin soil model. The latter model employs a variable
shear modulus and Poisson's ratio which are dependent on maximum shear
strain and hydrostatic pressure (1).
The purpose of this chapter is to discuss the implementation of a
new soil model into the CANDE program, called here the Duncan soil model,
and to present standard parameters for characterizing this model. The
Duncan soil model has had a substantial history of development and appli
cation over the last decade (26, 27, 28, 29 and associated references).
It is a variable modulus model such that increments of stress are re
lated to increments of strain by the isotropic form of Hooke's law wherein
the elastic paremeters are dependent on the stress state. For plane
strain, this incremental relationship may be written as:
Aa
x
Aa
y
At
11
12
C 12 °
Ae x
c 22
Ae y
o c„
ay
9.1
where Aa .,Aa ■ normal stress increments
x y
At = shear stress increments
Ae .,Ae  = normal strain increments
x y
Ay ■ shear strain increments
C ■ constitutive matrix components (variable)
In accordance with Hooke's law (isotropic form), the matrix com
ponents C . . are all defined with any two elastic parameters. Table 9.1
115
Table 9.1 Elastic Equivalents for
Isotropic Plane Strain
Matrix
Component
(E, v)
(E, B)
C ll  C 22
E(lv)
3B(3B + E)
9B  E
(1+v) (l2v)
C 12
Ev
3B(3B  E)
9B  E
(1+v) (l2v)
C 33
E
3BE
93E
2 (1+v)
116
shows this relationship for the elastic parameters pertinent to this
study: Young's modulus and Poisson's ratio (E, v) and Young's modulus
and bulk modulus (E, B) . If (E, v) or (E, B) are described as a function
stress, characterizing the nonlinear behavior of soil, then the matrix
components C.. are also defined and infer tangent relationships between
stress and strain increments.
9.1 DUNCAN MODEL REPRESENTATION OF ELASTIC PARAMETERS
Initially, Duncan and his colleagues characterized soil behavior
with a variable tangent Young's modulus E t and a constant Poisson's
ratio where Fj employed the socalled hyperbolic stressstrain model (27)
Subsequently, a variable Poisson's ratio formulation was introduced to
better represent the volume change behavior observed in triaxial soil
tests (26,28). Recently, a tangent bulk modulus formulation was intro
duced to replace the variable Poisson's ratio (29).
The last model, which employs tangent Young's modulus and tangent
bulk modulus formulations, is adopted for this study and incorporation
into CANDE. An extensive evaluation of the variable Poisson ratio
formulation was undertaken during the course of this study and was
found to behave erratically in some cases (30). Consequently, it was
not incorporated into the CANDE program.
Development details of the Duncan model are well documented else
where (29). The final expressions for tangent Young's modulus and
bulk modulus as a function maximum and minimum principle stresses for
loading conditions are given here.
'The tangent Young's modulus expression is:
a R (lsin<f>)(a a )
E = KP (^) [1  — —  X . J 9.2
t a P 2 c cos<}> + 2a sxn<{>
3. J
where aj = minimum principle stress (compression positive),
o"! = maximum principle stress (compression positive).
P a = atmospheric pressure (for dimensionless convenience) .
K = modulus number, nondimensional
117
n = modulus exponent, typical range 1.0 to 1.0.
Rf = failure ratio, typical range 0.5 to 0.9.
c = cohesion intercept, units same as P a .
<j> = friction angle, radians
A<J> = reduction in <J> for 10fold increase in 03.
(i.e., 4  <j) A(J» log ,_A
xu ^Pa'
The tangent bulk modulus expression is a function of minimum compressive
stress given by:
9.3
where K_ = bulk modulus number, dimensionless.
m = bulk modulus exponent, typical range 0.0 to 1.0.
In Equations 9.2 to 9.3, there are a total of eight parameters to
define a particular soil in loading: K, n, Rf, c, <£ , and A<J) to define
E t ; and K^ and m to define B t . Established methods for determining
these parameters from conventional triaxial tests have been reported by
Duncan and his colleagues (28,29). In the last section of this chapter,
conservative estimates of these parameters are given for various soil
types and degree of compaction.
The behavioral characteristics and limitations of the Duncan soil
model (Equations 9.2 and 9.3) are enumerated below along with the
programming strategy used in the Duncan finite element program called
SSTIPN.
(1) As 03 increases (e.g. confining pressure in a triaxial test)
E t and B t becomes stiffer (assuming m and n are greater than
zero). However, as maximum shear stress increases (i.e.
(a.  cO/2), E t becomes weaker, but B« remains constant. Such
1 £ u
behavior is typical of triaxial tests on which the model was
developed.
(2) Shear failure is said to occur when E t approaches zero. That
is, the bracketed term in Equation 9.2 approaches zero as
118
a  a increases. If a significant portion of the soil mass
fails in shear, the results may no longer be reliable because
the model is not applicable for soil instability. To avoid
numerical problems, the SSTIPN algorithm arbitrarily limits
the minimum value of the bracketed term in Equation 9.2 to
1  .95 Rf. Thus, Et does not actually become zero in shear
failure.
(3) Tension failure is said to occur when a 3 becomes tensile. In
such cases the soil stiffness breaks down and cannot carry load.
To cope with this problem, the SSTIPN algorithm computes a small
value for bulk modulus from Equation 9.3 by specifying a /P = 0.1
3 Si
whenever 03 is tensile. For the second elastic parameter,
Poisson's ratio is arbitrarily assigned the value 0.495 and
Equation 9.2 is ignored. This results in equivalent Young's
modulus whose value is approximately 3% of the bulk modulus.
(4) In addition to the special treatment for shear and tension
failures, the SSTIPN algorithm sets limits on B t as predicted
from Equation 9.3 dependent on the value of E t from Equation 9.2.
Specifically, B t = E t /3.0, if B t is less than this value, and,
Bt = 34.0 E t , if B t is greater than this value. These limits
correspond to maintaining the equivalent Poisson's ratio within
the range 0.0 to 0.495.
(5) For each load step (e.g. construction increment), the SSTIPN
algorithm utilizes two iterations to determine B^ and E t as
defined above. For the first iteration, the stresses existing
in the element at the end of the previous load step are used
to estimate B t and E t to obtain approximate stress increments.
The second iteration repeats this solution wherein B t and E t
are now determined by adding onehalf of the stress increments
determined in the first iteration to the previous stress state.
Upon completion of the second iteration, the stresses are accumu
lated and printed out, and the next load step is considered. No
convergence check is made.
119
(6) When an element first enters the system, the "existing" stress
state used to determine B t and E fc for the first iteration is
determined in a special manner depending on whether the element
is part of the initial system (e.g. preexisting foundation) or
part of a new construction increment. For the case of elements
belonging to the initial system, existing stresses are defined
by the user (input), or if the foundation is composed of hori
zontal rows, initial stresses can be automatically approximated
by overburden pressure and a lateral coefficient.
Elements belonging to a new construction increment do not
have an existing stress state prior to entering the system.
However in order to evaluate E t and B t for the first iteration,
the SSTIPN algorithm estimates initial stresses based on element
height, soil density, assumed Poisson's ratio, and humped surface
angle.
(7) According to published reports (28,29), "unloading" of the
Duncan soil model is accomplished by replacing the tangent
Young's modulus function (Equation 9.2) with an unloading
expression; E u = K P a , 3vn, where 1^ is an unloading modulus
( p )
number whose value is greater than K.
Although this is relatively easy to program, there are
serious theoretical objections to this description of unloading.
Presumably, the criterion for unloading (i.e. switching from
E t to E u ) is by observing a decrease in maximum shear stress
irrespective of 03. Such a criterion may be sufficient for
load paths where 03 is constant (e.g. triaxial test), however
for more general load paths, serious violations of the contin
uity principle can occur, i.e., two arbitrarily close load
paths should not result in dramatically different stressstrain
responses.
For this reason, the unloading function is not incorporated
into the CANDE program. Further research on unloading is war
ranted.
120
9.2 CANDE SOLUTION STRATEGY FOR DUNCAN MODEL
The CANDE algorithm for the Duncan soil model is contained in a
new subroutine called DUNCAN. Here the representation of E t and B t
(Equations 9.2 and 9.3), shear failure, and tension failure are treated
in a similar fashion to the SSTIPN algorithm discussed in the previous
section. However, there are some significant differences in the CANDE
solution strategy with regard to (a) number of iterations, (b) averaging
Et and B t over a load step, and (c) treatment of elements entering the
system for the first time. These differences are discussed below.
Iterations . As previously explained, the SSTIPN algorithm uses two
iterations per load step for all loading schedules. Preliminary studies
during this research indicated that using just two iterations can lead
to serious error in predicting E and B even when load increments are
relatively small (e.g. one layer of elements per construction increment).
To deal with this problem, the CANDE algorithm allows the maximum
number of iterations to be specified by the user. During the iteration
process, the current estimate of Et for each element is compared percen
tagewise with the previous estimate of E t . *If two succeeding estimates
of E t converge within a specified error tolerance for all elements, the
iteration process is terminated and algorithm advances to the next load
step. Should convergence not be achieved after the specified maximum
number of iterations, a warning message is printed out prior to advancing
to the next load step.
Note, the convergence check is only considered for E t , not B t .
However, it may be presumed that B t converges more rapidly than Et
since the former is only a function 03, whereas the latter is a more
sensitive function dependent on a~ and 03.
Averaging Et and B t . Equations 9.2 and 9.3 are tangent moduli expressions
for E t and B t for a particular principle stress state a^ and 03. As a
load increment is applied, the stress state changes, inferring changes in
Et and Bt. In order to adequately represent the effects of these changes
in Equation 9.1, E and B should represent "average" values over the
load step. This, of course, is the purpose of iteration.
121
One way of obtaining average values is to evaluate Et and B t based
on the average stress state during the load step as is done in the SSTIPN
algorithm. Alternatively, one may average E t at the beginning of the
load step with E^ at the end of the load step. Likewise for B t . Speci
fically, this may be written as:
E avg " ar^'+rtj 9.4
B avg " ( 1  r > B l H ' rB 2 9  5
where E , B = E , B at startofloadstep (known)
E ? , B = E , B at endofloadstep (iteratively determined)
r = averaging ratio, (generally r = 1/2)
For reasons to be subsequently discussed, the CANDE algorithm employs
the averaging scheme given by Equations 9.4 and 9.5. Comparison studies
between the stress averaging scheme and the moduli averaging scheme were
found to give nearly identical results for r = 1/2.
The averaging ratio r is treated as a material input parameter in
the CANDE program. Generally r = 1/2, however for preexisting soil
zones, r = 1 permits proper calculation of preexisting stresses as dis
cussed next.
Entering Elements . Soil elements enter the structural system in one of
two categories. The first category apply s to preexisting or insitu
soil elements in which an initial stress state exists but is unknown.
Elements entering in this category are part of the initial configuration
and belong to the first construction increment.
The second category applys to fill soil elements, i.e., soil layers
added to the system in a predefined construction schedule. Here, the
initial stress state is nonexistant prior to entry into the system.
Both categories present special starting problems for the iteration
procedures because the initial stress state is unknown or undefined.
If preexisting soil zones are to be characterized by the Duncan
soil model, the 'initial stress state can be determined iteratively by
122
assuming the preexisting soil zone is a construction increment loaded
with its own body weight (and, if desired, a consolidation pressure).
Here the averaging ratio should be set to 1.0, so that, E a and B a
are equal to the endofloadstep values E£ and B2, respectively, and
correspond to the existing stress state. Beginningofloadstep values
El and B^ are initially set to 0.0 when an element enters the system.
However when r = 1, they have no influence on the averaging process.
After the first construction increment is complete, the program auto
matically sets the value of r to 1/2, so that, all subsequent moduli
calculations represent load step averages.
Elements entering the system in the second category have no initial
stiffness prior to loading so that Ej_  B]_ = 0. Accordingly, using
r = 1/2 gives average moduli values equal to onehalf of the endofload
step values, E and B .
To start the iteration process for entering elements of either
category, some guess must be made for E2 and B2 in order to construct
the first trial stiffness matrix. This is achieved by arbitrarily de
fining "dummy" principle stresses from which initial estimates of E2
and B2 are calculated. The dummy principle stresses have no effect on
the final values of E2 and B2» however they do influence the number of
iterations for convergence. Once an element has entered the system, the
initial guess for E2 arid B2 for all subsequent load steps are equated
to the last calculated values, thus dummy stresses are not required.
Cande Algorithm . Figure 9.1 is a flow chart of the CANDE algorithm
illustrating the solution strategy previously described. Some of the
limit bounds are defined differently than in the SSTIPN algorithm. For
example, the maximum equivalent Poisson ratio, v max , is set at 0.48
rather than 0.495 in order to avoid unreasonably high values of C in
Equation 9.1. Also, the shear failure factor (1D) is assigned a lower
limit of 0.05, rather than 10.95 Rf in order to provide a greater re
duction of stiffness in shear failure.
123
Incoming information,
a , a  principle stresses
i = iteration no.
i
■'/
Existing element
New element
E =
B i = °
a, = 0.2 P
1 a
E 1 = E 2
B 1 = B 2
r = 0.5
i = 1
i > 1
>
a 3 = 0.1 P a
.._ „. }
1
Check tension failure.
a 3 <
t w I
yes
i
no
<
*
)
Set lower limit on a .
Set tension values
a > 0.1 P
3  a
E 2 = (0.05) 2 K(.l) n
B 2 = 1.67E
*
f
Set limit on shear failure.
D = R f ( a 1 " a 3 ) sin< J )
2 (c cos<j) + a sin<j>)
< D <_ 0.95
'
t
Compute endof'step moduli.
E, = K(a./P ) n (lD) 2
i j a.
B 2 = VW
E 2 /3 £ B 2 < 8E 2
'
f
<k
1
*
Average moduli values.
E = (lr)E. + rE.
avg 1 2
B = (lr)B n + rB„
avg 1 2
.... . _ .. 1
*
fl ^y Re P eat iteration.
Check convergence of E„.
E 2 (i) = E 2 (i+1) ?
"^si*
c
A Go
to next loac
[ step.
Figure 9.1 CANDE Algorithm for Duncan Soil Model.
124
Two additional features of the CANDE algorithm not shown in Figure 9.1
are; (1) an under relaxation scheme to improve the rate of convergence
for E2, and (2) a constant Poisson ratio option which replaces the tangent
bulk modulus formulation.
The under relaxation scheme comes into play after the second iteration
wherein each estimate of E£ is a weighted average of the current estimate
and the previous estimate. This feature takes advantage of the observation
that E2 generally converges in an oscillatory manner.
When the constant Poisson ratio option is exercised, all references
to the bulk modulus formulation are bypassed. Otherwise, the algorithm
is essentially the same.
9.3 STANDARD HYPERBOLIC PARAMETERS
Whenever possible, the hyperbolic parameters characterizing the
Duncan soil model should be determined directly from triaxial tests
using established curvefitting procedures (28,29). In many instances,
however, triaxial data may be unavailable, and so, it is convenient to
establish "standard" parameter values for various types of soil and
degrees of compaction. Table 9.2 (abstracted from Reference 29) provides
parameter values for four soil classifications, each with three levels
of compaction. These "standard" values are conservative in the sense
that they are typical of lower values of strength and moduli observed
from numerous triaxial tests for each soil type. An independent study (30)
to establish standard parameters for E t utilizing the same data base is
in good agreement with Table 9.2.
For convenience, the hyperbolic parameters in Table 9.2 are stored
in CANDE and may be used by simply identifying soil type and level of
compaction.
The behavior of the Duncan soil model for simulated uniaxial strain
and triaxial loading tests is shown in Figures 9.2 through 9.5 for the
standard parameters in Table 9.2. Specifically, Figure 9.2a shows axial
stress vs. axial strain in confined compression for three compaction
levels of coarse aggregates. The slope of these curves is the tangent
125
CO
5i
CD
41
CD
6
CO
5i
CO
Pi
O
X>
5i
0)
a
T3
51
CO
§
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CM
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B
CN CN CM
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• • •
o o o
m m m
dod
CM CM CM
• • •
o o o
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m m o
r*» r^ m
H
o o o
m m m
<r cm iH
o m o
o r>» m
CM
o o o
<t co m
rH
141
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r — r~^ r^
CD CD o
r — r^* r— ~
d o d
r>. f*. r»
• • •
o o o
r» r>« r.
odd
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sr a d
• • •
o o o
m m m
CN CM CM
• • •
o o o
vO vO vO
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si <■ a
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41
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ci in co
co «<r cm
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41
e £
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J*
0.150
0.140
0.135
0.135
0.125
0.120
0.135
0.125
0.120
0.135
0.125
0.120
RC
Stand.
AASHTO
m m o
O Q\ o\
iH
o o m
O C\ CO
o o m
O CJ\ 00
rH
o o m
o o\ CO
rH
Unified
Classification
Coarse Aggregates
GW, GP
SW, SP
Silty Sand
SM
Silty Clayey Sand
SMSC
Silty Clay
CL
CO
6 CM
rs
S3
2
rH <T
II II
CO CM
41 4J
■41 <4(
a a,
•H H
c
o
•H
41
CJ
(0
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8
41
•H
o
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CJ
c
<D
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T3
>
•H
41
41
,fl
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00
rH
•H
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pej
rs
II
ii
CJ
6
PS
>
126
confined modulus (i.e. Cq in Equation 9.1) and are observed to increase
with axial stress and compaction level. Figure 9.2b shows the behavior
of the same soil models in triaxial loading. Here, the slope of the
curve is the tangent Young's modulus E t and are observed to decrease with
shear stress and increase with compaction level as expected.
The remaining three pairs of figures illustrate the same trends for
other soil types. The siltysand soil type (Figures 9.3a,b) have the
largest stiffness values (slopes) in confined compression, while the
siltyclay (Figures 9.5a,b) have the lowest.
127
RC = 90
Figure 9.2a.
8 9 10
Percent axial strain
Coarse Aggregates, Uniaxial Behavior
12
tfft
 RC=105
5 6 7 8 9 10 11 12
Percent axial strain
Figure 9.2b. Coarse Aggregates, Triaxial Behavior
128
Pa
RCzlOO RC=90 RCi85
£
4
(/)
x
<
■» £■
7 8 9 10 11 12
Percent axial strain
Figure 9.3a Silty Sand, Uniaxial Behavior
gjft
7 8 9 10 11 12
Percent axial strain
Figure 9.3b Silty Sand, Triaxial Behavior
129
RC=90
RC=85
7 8 9 10 11 12
Percent axial strain
Figure 9.4a. Silty Clayey Sand, Uniaxial Behavior
<fl 10
RC100
Figure 9.4b,
7 8 9 10
Percent axial strain
Silty Clayey Sand, Triaxial Behavior
130
RC= 85
>6i
Percent axial strain
Figure 9.5a. Silty Clay, Uniaxial Behavior
o;(r 3
Percent axial strain
Figure 9.5b. Silty Clay, Triaxial Behavior
131
CHAPTER 10
SUMMARY AND CONCLUSIONS
This report presented a step by step development for the structural
analysis of buried, precast reinforced concrete box culverts using the
finite element method to model the soilbox system. Model predictions
were validated with measured data from both inground and outofground
experimental tests. A user oriented soilbox model with automated finite
element mesh generation is operational in the CANDE1980 computer program
and is referred to as "level 2 box" option. Also operational in CANDE1980
is the Duncan soil model with simplified input options for standard types
of soil. Specific findings and conclusions from this work are listed below,
1. Loaddeformation curve predictions for reinforced concrete
culverts are sensitive to the cracking strain parameter e .
However at ultimate flexural capacity, the maximum stresses of
concrete f and steel f y are the controlling parameters with
regard to the reinforced concrete model.
2. CANDE predictions are in good agreement with experimental data
from reinforced concrete pipes in threeedge bearing. Better
correlation for loaddeformation curves was observed for pipes
failing by flexure than by shear. Predicted ultimate loads,
whether in shear or flexure, are within 10% of measured values.
3. The measured cracking load and ultimate load for reinforced
concrete box culverts tested in fouredge bearing show good
correlation with CANDE predictions. Predicted cracking loads
averaged 10% lower and ultimate loads averaged 1% lower than
experimental results. Predictions from the SGH design/analysis
approach are similar to CANDE predictions but showed slightly
more deviation from experimental data.
132
A. Measured soil pressures from a full scale burled box installation
are in good agreement with CANDE predictions at both intermediate
and final burial depths. Vertical soil pressures on top and
bottom slabs were in very close agreement, whereas measured
lateral pressures on the sides at full burial depth were some
what lower than predicted.
5. Assumed soil load distributions on buried boxes used in the
development of ASTM C789 design tables were compared with pre
dicted soil load distributions determined from CANDE resulting
in the following observations. (a) Vertical soil pressure on
the top and bottom slabs are not uniform as assumed but increases
monitonically from the centerline to the corners. (b) Shear
traction on the sidewalls produces a significant downward force
that must be equilibrated by an upward pressure on the bottom slab.
(c) Soil stiffness is an important parameter for determining soil
load distributions and magnitudes. The latter two effects are not
presently taken into account in the ASTM loading assumptions.
6. Based on CANDE predictions, the design earth covers specified in
ASTM C789 design tables are generally conservative providing
good quality backfill soil is assumed. However, specified earth
covers for boxes with high levels of reinforcement tend to be
less conservative than specified earth covers for boxes with low
levels of reinforcement.
133
APPENDIX A
DETAILS OF REINFORCED CONCRETE MODEL
The reinforced concrete beamrod model presented in Chapter 3 is
discussed in further detail in this appendix. This model replaces the
original CANDE concrete pipe type and can be used with solution levels
1, 2 or 3. ; Subroutine CONMAT is the heart of the new reinforced concrete
model wherein concrete cracking, loading to ultimate and unloading is
simulated.  For purposes of this appendix, it is presumed the reader is
familiar with basic assumptions and general solution strategy presented
in Chapter 3. Here attention is focused on programming details.
I. VARIABLES USED
To calculate the initial and load dependent mechanical properties
of the reinforced concrete sections, parameters describing the material
behavior of concrete and steel have to be defined. Some of these para
meters are primary (defined by input) while others are secondary (de
rived from primary) . The main purpose of these parameters is to define
an idealized stressstrain diagram for concrete and for the steel rein
forcement (see Figure A. 1, A. 2) . The following parameters are primary
input data for the material properties, where, in parentheses, are the
default values used in CANDE.
e = concrete strain at tensile cracking (0.000 in/in)
e = concrete strain at elastic limit (0.5 f'/E, )
y c 1
e' = concrete strain at f (0.002 in/in)
c c
f = unconfined compressive strength of concrete (4000 psi)
c
E = Young's modulus for linear concrete (33 vf 1 (y) " )
1 c c
134
Strain
Cracking
Figure A.l  Idealized Concrete StressStrain Diagram.
00
0/
%
/
5/
f
/
/
7
/
#
Strain
Figure A. 2  Idealized Steel StressStrain Diagram.
135
v = Poisson's ratio for concrete (0.17)
c
Y = unit weight of concrete (150 pcf)
c
f = yield stress of steel (40000 psi)
E = Young's modulus of steel (29 x 10 psi)
o
v = Poisson's ratio of steel (0.30)
s
The following are secondary parameters derived from primary data:
2
E = confined elastic modulus of concrete (E /(1v )
c 1 c
v = shear strength of concrete (2/ f ' psi)
c c
f = maximum tensile stress of concrete (e x E )
f = concrete stress at elastic limit (e x E., )
yc y 1
2
E = confined elastic modulus of steel (E /(1v )
s os
n = concretetosteel modulus ratio (E /E )
c s
In addition, the analysis mode requires (see Figure A. 4):
h = wall thickness of concrete (in)
As. = area of inner reinforcement per unit length of pipe
(in /in)
2
As^ = area of outer reinforcement per unit length of pipe
(in /in)
2
c. = concrete cover on inner reinforcement (1.25 in)
c = concrete cover on outer reinforcement (1.25 in)
o
Using these parameters the initial uncracked section properties are
defined as:
Effective axial stiffness:
EA* = E (h + (n1) As. + (n1) As )
C X o
136
Neutral axis of bending:
2
_ (7 + As.(nl)c. + As (nl)(hc ))
y = E 2 x 1 o o
c _
EA
Effective bending stiffness:
3
EI* = E [yr + (£  y) 2 h + (n1) (As.(c.y) 2 + A (hcy) 2 ]
c LI I 11 so o
The above section properties EA , y and EI are for the uncracked
cross section of 1 inch width with no loading. Now that the initial
stiffnesses of the beam rod sections are defined, the beamrod elements
are ready to be analyzed for the first load increment.
II. PROCEDURE TO CALCULATE THE SECTION PROPERTIES
Using the initial section properties above, the structural system
is solved for the first load increment resulting in trial solutions for
thrust and moment increments within each element. Due to the nonlinearity
of the materials, the initially assumed section properties are modified
and another trial solution is obtained. This iterative solution technique,
which considers the average stress state during the increment to find
effective section properties, is repeated until convergence within the
load step is achieved.
CONMAT subroutine evaluates EA , y, and EI for each element until
all elements converge as the system advances from load step i1 to load
step i. The procedure used for each load step is as follows:
a) For each element an increment of moment and thrust is obtained
from the general solution process. If a section initially cracks or
extends its crack, the stresses in the newly cracked region are zero,
thus, the preexisting stresses prior to cracking must be redistributed.
This redistribution can be achieved with corrections to the thrust and
moment increment (called here thrust and moment redistribution). In
our approach this correction is made after the inner loop convergences,
so that, initially, moment and thrust redistribution is zero (see step h
137
for redistribution).
* _ *
b) Assuming the section properties EA , y, and EI from the con
verged solution at load increment i1 and using the increment of moment
and thrust, a linear strain distribution for the section is calculated
(see Figure A. 3).
, AN , AM , • '
e i = e ii + m* + ei* (y " y)
where
e . 1 = strain distribution from converged solution at load
increment i1
AN = increment of thrust + (thrust redistribution, last
iteration)
AM = increment of moment + (moment redistribution, last
iteration)
y = spatial coordinate from section bottom.
This linear strain distribution e. is the first tentative solution
1
of the iterative procedure.
For computational convenience, the section stiffness properties of
EA and EI (defined in Chapter 3) are divided by the confinedelastic
* *
concrete modulus E , so that we may define A and I as:
c
it *
A = EA /E
* * c
I = EI /E
c
Or more explicity, to obtain section properties the following integrals
must be evaluated (see Figure A. 4):
A = FE(y) dy + WSI (n1) A +(WSO)(nl) A gQ
' o
rh
o
FE(y) y dy + WSI (n1) y . A . + WSO (n1) y fc A
J ' J * ti si to so
A
I* = FE(y)(yy) 2 dy + WSI(nl) (y^y)^. + WSO (n1 ) (y^?)^
138
sh A £ n
n
6rl Si
+
Figure A. 3  Strain Distribution at Load Increment i.
•AS,
»AS;
«t,
yti
4— X^4
Figure A. 4  Integration Points of the Cross Section.
139
where:
FE(y) = E" /E (modulus reduction ratio for concrete)
c c
WSI = E' /E (for inner reinforcement reduction)
WSO  E' /E (for outer reinforcement reduction)
s s ■
E'  tangent modulus of concrete
c
E' = tangent modulus of steel
s
Using the above formulas and the strain distribution, e . , the new
properties EA , y and EI are calculated. From the formulas it is ob
served that the concrete part of the section requires an integration of
FE(y) to evaluate its properties, where the function FE(y) is not smooth.
Thus, Simpson's integration is performed for the concrete section using
eleven points along its depth (see Figure A. 4).
c) The concrete is analyzed first using the strain distribution
e . _ and £.. The integration points are analyzed one at a time, dependent,
in part, on the value FE(y). The first step is to update the record for
the maximum stressstrain occurrence of each point during its loading
history. Specifically, the converged strain distribution of load step
i1 and the maximum strainstress values of the point computed in previous
steps are used to identify if there are new maximum values to be saved.
If the strain of the point at load step i1 is less than any previous
maximum strain, no change is made in the history vector. If the strain
of the point at load step i1 is greater than the maximum strain, a new
maximum was reached for that point at load step i1. Using the new
maximum strain, the new maximum stress is computed using the old maximum
stressstrain value located along the basic stressstrain diagram (see
Figure A.l). The value of the maximum stressstrain of the point is
saved and subsequently used to define the unloading and reloading path
of the stressstrain diagram at load step i.
d) A comparison of the strain e . with e . 1 for the integration
point indicates if the point is loading or unloading. If the point is
unloading ( e <  e J ) the stress a of the point can be calculated
140
knowing that the unloading is elastic,
a . = o + (e.  e )E
1 max 1 max c
where (a , e ) = maximum stressstrain value
max max
e . = strain of the point at load step i
If the point was not previously cracked and a. doesn't reach the
concrete tensile strength then FE(y) = 1.0, if the point was previously
cracked and a. is in compression then FE(y) = 1.0. Otherwise FE(y) is
0.0 (see Figure A. 5).
If the point is loaded, it can be initial loading or reloading.
Using the previous strainstress values (e . , a. ) and the maximum
strainstress of the point (e , a ) , the loading or reloading cases
max max
are identified. If the previous values are equal to the maximum it
means initial loading, otherwise it means reloading. Once the case is
defined and the values of (e . ,, a. ,) and (e , a ) are known, the
x1 ll max max
case falls in one of the possible point histories presented in Figure A. 5.
Knowing in what case the point is, the stress value a. of the point for
load step i is evaluated. Now that the stressstrain value (a., e.) is
11
defined. E' is determined by the slope between (e . n , a. .) and (e . , a.).
c li 11 1 1
Thus FE(y) for each point is:
1 11 c
This procedure is executed until FE(y) is defined for the eleven inte
gration points.
e) In the previous step (d), when the value of FE(y) is determined
for each integration point, at the same time, it is possible to determine
what points are cracked or uncracked. With this data and doing a linear
interpolation of stresses between the points of crack and no crack, the
crack depth of the section is calculated. This calculated crack depth
is printed out. Note crack depth is completely different than crack
width discussed in Chapter 3.
141
7
>S
* (e^* cr^) load step i1  FE(I)
X ^ C i* °i^ load step i FE(I)
Ei I
s tangent modulus i
\= maximu m strainstress history
E (Z^
0.0 (vhen cracked and
a. less than zero)
Figure A. 5  Modulus Function for all Possible Concrete Strain Histories
at a Point in the Beam Cross Section.
 142
f) Knowing the cross section strain distribution £•_]» e  an ^
the location of the reinforcement in the cross section, the strains of
the inner and outer reinforcement are calculated for load step i1 and
i. Using the stressstrain values at load step i1 and the strain value
at load step i for the reinforcement, the value of the stress at load
step i is calculated for elastic loading or unloading as:
o . = a .  + (e  e . ) E
i ll 1 iI s
For plastic loading we have:
a. = f (steel yielding stress)
This procedure applies to both inner and outer reinforcement. Using
their stressstrain values, the factors WSI and WSO for the inner and
outer reinforcement are calculated respectively as follows :
(a i ~ a il }
WSI = , r (stressstrain values for the inner
< e i " E i1> E s
reinforcement )
( °i " ? il )
WSO = — ( r (stressstrain values for the outer
i il s reinforcement)
All the possible cases of strain histories for the steel reinforcement
are shown in Figure A. 6.
g) With the preceeding developments, the section properties defined
in step b are evaluated with the aid of eleven point Simpson integration
as follows:
A = SUM1 + SI + SO
 = (SUM2 + y. SI + y SO)/A
y xo
i* = y 2 sumi  2y sum2 + SUM3 + (y^y) 2 si + (y Q y) SO
In the above, SUMI, SUM2, and SUM3 represent the concrete contri
butions and are the numerical integrations of the integrands FE(y),
2
yFE(y) and y FE(y), respectively, i.e.,
SUMI  4? (FE(1) + 4FE(2) + 2FE(3) + ••• FE(ll))
143
(£. ,, "O^i) l° a <i step i1
X (Sj, <7.) load step i
Ee= steel Young's modulus
E* = tangent modulus
wsi(o)  eVe_
Figure A. 6  Modulus Function for all Possible Steel Strain Histories
at a Point in the Beam Cross Section.
144
SUM2 = 4^ (FE(l)y + 4FE(2)y 2 + 2FE(3)y 3 + ••• FE(ll)y )
SUM3 = 4f (FE(l)y? + 4FE(2)yJ + 2FE(3)y^ + ••• FE(ll)y^)
Steel contributions SI and SO are associated with inner and outer
reinforcement and are given by:
SI = WSI(nl)A .
si
SO = WSO(nl)A
so
If either or both reinforcements are located in a cracked zone, the value
of "n" is used instead of "n1". That is, n1 accounts for the reinforce
ment hole in uncracked concrete.
* _ *
These values of A , y and I are compared with the assumed values
at step (b). The value A is used to check convergence. If the con
vergence is not reached, a new set of strains e. are evaluated at step
(b) using the new A , y, I . The entire procedure is repeated until
•A* mm J
successive values of A , y, and I converge or after four iterations
(inner loop), then the program goes to the next step h.
h) Now that new values of A , y and I are known for the cross
section at load step i (however, convergence of AM and AN are not yet
assured), a modified set of strainstress values (e. , a.) for each
point in the cross section is computed to account for stress redistri
bution due to cracking. These values are compared with the strain
stress value (e. ,, a. ,). If for some point the stress a. implies
ll ll i
that the point is cracked and was not previously cracked, it means
that the stress a. must be redistributed to the remaining uncracked
concrete and reinforcement. The procedure adopted in the program is
to evaluate an equivalent moment and thrust to be redistributed due to
the cracking of those points at load step i, where:
AN = Z Ay • a (thrust redistribution)
AM^ = Z Ay * o (yy) (moment redistribution)
145
After the moment and thrust redistribution is complete the program is
ready to go to step (a) and do the same procedure for the next element
(node) .
i) All elements of the reinforced concrete structure are analyzed
4c — 4c
and the values of A , y, I , AN R and AM are calculated for each one.
4c ~
If the values of A and y obtained for every element converge with the
ones obtained in the previous iteration and AN , AM are zero for all
R R
the nodes, the trial solution has converged. Otherwise, the system is
* °" 4c
solved again using the values of A , y and I from the last iteration.
This is called "inner loop" iteration. Once the inner loop converges,
the section properties, EA , y, EI , are used to get another solution to
the entire soilstructure system. This gives new values for AM and AN
to repeat the inner loop. Successive solutions for AM and AN is called
outer loop iteration.
Outer loop iteration is repeated until a convergence is reached
between successive solutions for AM and AN. If convergence is not
reached after six trial response solutions, the program assumes the
last one as an approximate solution and advances to next load step.
146
APPENDIX B
CANDE1980: USER MANUAL SUPPLEMENT
The following user's guide is a supplement to the 1976 "CANDE USER
MANUAL" (2). This supplement provides input instructions for the new
options on reinforced concrete box culverts and Duncan soil model described
in the main body of this report.
The original 1976 manual is still the principal reference source and
may be used without reference to this supplement if the new options are
not desired. Taken together, the original and new options form the program
called CANDE1980. Input instructions for CANDE1980 follows the same
pattern as the original program, composed of three main sections (A,B, and
C) as shown in Figure B.l, where new options are marked with an asterisk.
Section A is the master control input (card 1A) and is unchanged from
the original program. Section B (cards IB to 3B) includes a new input
option for modeling reinforced "concrete box" culverts in addition to
the original pipe types. The "concrete box" culvert type is only operative
in the analysis mode and cannot be used with solution level 1. Section C
(card sets C and D) includes the new "level 2 box" finite element generation
scheme discussed in Chapter 6 along with the new Duncan soil model option
presented in Chapter 9.
The supplemental input instructions to be given here provide a complete
set of data input for the subset of options shown in Figure B.2. Thus,
these instructions are self contained for "level 2 box" solutions with any
soil model. For extended level 2 and level 3 options, however, the 1976
user manual must also be used where noted.
Formatted input instructions for Sections A, B, and C are presented
in order, followed by explanatory comments and illustrations. Example
inputoutput data is given in the next appendix.
147
SECTION A  MASTER CONTROL: CARD 1A
• Execution mode = design or analysis
• Culvert (pipe) type ■ steel, etc.
• Solution Level = 1,2, or 3 (option, level 2extended)
SECTION B  CULVERT TYPE INPUT: CARDS IB up to 3B
Steel
1B,2B
Aluminum
1B,2B
Concrete
Pipe
1B,2B,3B
Plastic
1B,2B
Basic
1B,2B
Concrete
box
1B,2B,3B
SECTION C  SOLUTION LEVEL INPUT: CARDS CI up to C7
Level1
1C,2C
Level 2
pipe
1C,2C,3C
Level2
box*
1C,2C
Level2
extended
4C to 7C
Level3
1C to 5C
"V
(End)
i
SOIL MODELS: CARDS Dl up to D4
Elastic
ID, 2D
Ortho
Elastic
ID, 2D
it
Duncan
model
ID to 4D
Over
Burden
ID, 2D
Hardin
model
ID, 2D
Inter
face
ID, 2D
* New CANDE Options.
FIGURE B.l General Input Flow for CANDE1980
148
SECTION A  MASTER CONTROL: CARD 1A
• Execution mode = analysis
• Culvert type = concrete (box)
• Solution level = 2
SECTION B  CONCRETE BOX INPUT: CARDS IB to 3B
• Reinforced concrete material properties (IB, 2B)
• Option for standard crosssections (3B1, 3B2)
• Option for arbitrary crosssections (3B)
SECTION C  LEVEL 2 BOX: CARDS 1C, 2C, ID to 4D
• Installation type (1C)
• Boxsoil dimensions (2C)
• Soil model type, density (ID)
 Elastic (2D)
 Orthotropic elastic (2D)
 Duncan model (2D to 4D)
 Overburden dependent (2D)
 Hardin model (2D)
 Interface model (2D)
FIGURE B.2
Specific input flow for level2 box culverts
described in supplemental user's manual.
149
SECTION A  MASTER CONTROL CARD
Input
Card 1A. Master control card (one card per problem) :
Columns
(format)
0106
(A4,2X)
0808
(ID
Variable
(units)
XMODE
(word)
LEVEL
Entry Description
Word defining program mode,
= ANALYS, denotes analysis problem
= STOP, program terminates, last
card in deck
Defines solution level to be used
= 2, denotes finite element
solution with automated mesh
= 3, denotes finite element
solution with userdefined mesh
Notes
(1)
(2)
1015
PTYPE
(A4,2X)
(word)
1776
HED
(15A4)
(words)
7778
NPMAT
(12)
7980
NPPT
(12)
Defines pipe material to be used,
= CONCRE, denotes reinforced
concrete (pipe or box)
User defined heading of problem to be
printed with output
Number of pipe elements; only required
when LEVEL = 3
Number of pipe nodes; only required
when LEVEL = 3
(3)
*** GO TO SECTION B ***
150
SECTION B  BOX CULVERT
Reinforced Concrete Input (Cards IB, 2B, 3B)
Card IB. Material properties of reinforced concrete.
Columns
Variable
(format)
(units)
0110
PDIA
(F10.0)
(in.)
1120
PT
(F10.0)
(in.)
2225
RSHAPE
(A4)
(word)
Entry Description
Notes
2630
(15)
NONLIN
3140
STNMAT
(1)
(F10.0)
in/in.
4150
STNMAT
(2)
(F10.0)
in. /in.
5160
STNMAT
(3)
(F10.0)
in/in.
Any negative value. This signals (4)
program that is working with box culverts
Nominal concrete wall thickness. This (5)
value is used whenever wall thickness
is not specified on Card 3B
Control word to. select manner of (6)
input for cross section properties
on Card 3B
= ARBI, implies section properties
at each node along connected
sequence will be specified by
user (arbitrary)
= STD, implies simplified property
input will be allowed in conjunction
with Level 2 box mesh
Degree of nonlinearity , (7)
■ 1, concrete cracking only
= 2, also include nonlinear
compression of concrete
= 3, also include steel yielding
Concrete strain at which tensile (8)
cracking occurs (positive),
Default =0.0 in/in.
Concrete strain at elastic limit
in compression (positive)
Default = 1/2 PFPC/PCE (see next card)
Concrete strain at initial
compressive strength, f^, (positive),
Default = 0.002 in/in.
151
Card 2B.
Concrete and
Columns
Variable
(format)
(units)
0110
PFPC
(F10.0)
(psi)
1120
PCE
(F10.0)
(psi)
2130
PNU
(F10.0)
3140
PDEN
(F10.0)
(pcf)
4150
PFSY
(F10.0)
(psi)
5160
PSE
(F10.0)
(psi)
6170
PSNU
(F10.0)
7180
SL
(F10.0)
(in.)
steel properties:
Entry Description
Compressive strength of concrete, f£,
Default  4,000 psi
Young's modulus of concrete in
elastic range, ~.~ . , ?
Default = 33 (density) ' (f) '
c
Poisson ratio of concrete,
Default = 0.17
Unit weight of concrete (density)
Default = 0.0 for body weight;
however,
Default = 150 pcf for modulus
calculation
Yield stress of reinforcing steel,
Default = 40,000 psi
Young's modulus of steel
Default = 29 x 10 6 psi
Poisson' s ratio of steel,
Default =0.3
Spacing of reinforcement
Default = 2.0 in.
Notes
(8)
(9)
(10)
152
Card 3B. For RSHAPE = STD (only), two cards are required.
Only for Level 2 Box.
Card 3B1.
Concrete
Boxwall dimensions
Columns
Variable
Entry Description
0110
PTT
Thickness of top slab
(F10.0)
(in.)
Default = PT
1120
PTS
Thickness of side slab
(F10.0)
(in.)
Default = PT
2130
PTB
Thickness of bottom slab
(F10.0)
(in.)
Default  PT
3140
HH
Horizontal haunch dimens
(F10.0)
(in.)
4150
HV
Vertical haunch dimensio
(FIO.O)
(in.)
(Notes)
(11)
(12)
Card 3B2.
Steel rei
Columns
Variable
0110
(F10.0)
AS1
(in.2/in.)
1120
(F10.0)
AS 2
(in.2/in.)
2130
(F10.0)
AS 3
(in.2/in.)
3140
(F10.0)
AS4
(in. 2/in. )
4150
(F10.0)
XL1
5160
(F10.0)
TC
(in.)
Entry Description
Outer steel area, side wall
Inner steel area, top slab
Inner steel area, bottom slab
Inner steel area, side wall
Length ratio of AS1 steel along
top (bottom) slab
Uniform thickness of cover to all
steel centers
Default = 1.25 in.
(13)
*** GO TO SECTION C ***
153
Card 3B. For RSHAPE = ARBI (only) repeat this card for
number of pipe nodes (NPPT*) . This card may
be used for Level 2 or Level 3.
Entry Description Notes
Area of inner steel reinforcement (14)
Area of outer steel reinforcement
Thickness of concrete cover to
center of inner steel
Default = 1.25 in.
Thickness of concrete cover to
center of outer steel
Default = 1.25 in.
Thickness of concrete
Default = PT
* In Level 2 solution NPPT = 15
*** GO TO SECTION C ***
Columns
(format)
Variable
(units)
0110
(F10.0)
AS I
(in.2/in. )
1120
(F10.0)
ASO
(in. 2/in. )
2130
(F10.0)
TBI
(in.)
3140
(F10.0)
TBO
(in.)
4150
(F10.0)
PTV
(in.)
154
Columns
Variable
(format)
(units)
0104
WORD
(A4)
SECTION C  SOLUTION LEVEL DESCRIPTION
Level 2 Input (Cards 1C, 2C, ID, 2D)
Card 1C. Define mesh type, title, and special options:
Entry Description
Name to identify type of automatic
mesh
= EMBA, embankment mesh
 TREN, trench mesh
User description of mesh to be
printed with output
Command to permit user to selectively
modify the automatic mesh,
= MOD, mesh will be modified
^ MOD, mesh will not be modified
(left, justified)
0572
(17A4)
7376
(A4)
TITLE
W0RD1
Notes
(15)
(16)
For level 3 input, see Section C in 1976 manual.
155
Card 2C. Define print options and mesh parameters:
Columns
(format)
0105
(15)
0610
(15)
1115
(15)
1620
(15)
Variable
(units)
IPLOT
IWRT
MGENPR
NINC
2130
Rl
(F10.0)
(in.)
3140
R2
(F10.0)
(in.)
4150
HTCOVR
(F10.0)
(ft.)
5160
DENSTY
(F10.0)
(pcf)
Entry Description Notes
Signal to create a plot data tape
on unit 10
= 0, no data tape created
= 1, create data tape
Signal to print out soil response
for all elements,
= 0, no soil response printed out
= 1, print out soil response
Code to control amount of print out
of mesh data,
= 1, minimal printout; just
control data
= 2, above, plus node and element
input data
= 3, above, plus generated mesh data
= 4, maximal printout of input data
Default = 3
Number of construction increments, (17)
= 1, combine all lifts into
one monolith
= 0, used for data check only;
all data is read but not executed
= N, number of construction
increments to be executed,
N = 1 to 20
Distance from center of the box to (18)
center of side wall
Half of the distance from center of top
slab to center of bottom slab.
Height of soil cover over the top (19)
of the box
Density of soil above truncated mesh
to be used as equivalent overburden
pressure
continued
156
Card 2C.
continued
Columns
Variable
(format)
(units)
6170
TRWID
(F10.0)
(ft.)
7180
BDEPTH
(F10.0)
(in.)
Entry Description
Width of trench; only required
for WORD = TREN
Depth of bedding material
Default = 12 in.
Notes
(20)
(21)
For extended level 2 option (W0RD1=M0D) , insert cards 4C to 7C here prior
to card set D (See note 16). Otherwise, go directly to card set D.
157
Soil Data Cards:
Card ID. Material identifier card (repeat D cards for each material)
Columns
(format)
0101
(Al)
0205
(14)
0610
(15)
1120
(F10.0)
2140
(5A4)
Variable
(units)
LIMIT
ITYP
DEN (I)
(pcf)
MATNAM
(words)
Entry Description Notes
Last material cardset indicator;
= 0, read another set of material
definitions
= L, this is the last material
input
Material zone identification number (22)
for level 2 box use:
= 1, for insitu soil zones
= 2, for bedding zones
= 3, for fill soil zones
Selection of material model to be (23)
associated with material zone I,
= 1, linear elastic (isotropic)
= 2, linear elastic (orthotropic)
= 3, Duncan soil model
= 4, overburden dependent model
= 5, Hardin soil model
= 6, frictional interface
(not operative with Level 2 box)
Density of material I used to compute
gravity loads; not applicable for
ITYP = 6.
For ITYP = 3,4, or 5, MATNAM is used (24)
to select soil subgroup models as shown
in Table B.l on the next page. In all
cases, MATNAM is printed out with the
data but has no control for ITYP = 1, 2
or 6.
Go to card 2D corresponding to ITYP.
158
TABLE B.l. Soil models controlled by MATNAM (Card ID)
MATNAM
Soil Model Description
ITYP = 3, Duncan soil model, Chapter 9, Table 9.2
CA105
CA95
CA90
SM100
SM90
SM85
SC100
SC90
SC85
CL100
CL90
CL85
USER
coarse aggregate, relative compaction 105%
coarse aggregate, relative compaction 95%
coarse aggregate, relative compaction 90%
silty sand, relative compaction 100%
silty sand, relative compaction 90%
silty sand, relative compaction 85%
silty clayey sand, relative compaction 100%
silty clayey sand, relative compaction 90%
silty clayey sand, relative compaction 85%
clay, relative compaction 100%
clay, relative compaction 90%
clay, relative compaction 85%
Parameters supplied by user
ITYP = 4, Overburden Dependent, 1976 CANDE manual, pg. 39
GGOOD
GFAIR
MGOOD
MFAIR
CGOOD
CFAIR
USER
granular soil, good compaction
granular soil, fair compaction
mixed soil, good compaction
mixed soil, fair compaction
cohesive soil, good compaction
cohesive soil, fair compaction
parameters supplied by user
ITYP = 5, Hardin soil model, 1976 CANDE manual
GRAN
MIXED
COHE
TRIA
granular soil, specified void ratio
mixed soil, specified void ratio
cohesive soil, specified void ratio
parameters specified by user (triaxial test)
* MATNAM must be left justified (i.e. start in column 21)
Defaults are: MATNAM = USER for ITYP = 3 and 4, or
MATNAM = MIXED for ITYP = 5.
159
Card 2D. ITYP = 1, linear elastic
Columns
(format)
0110
(F10.0)
1120
(F10.0)
Variable
(units)
E
(psi)
GNU
Entry Description
Young's modulus of materil I
Poisson's ratio of material I
Notes
(25)
Card 2D, ITYP = 2, orthotropic, linear elastic:
Columns Variable
(format) (units) Entry Description
0110 CP(1,1) Constitutive parameter at matrix
(F10.0) (psi) position (1,1)
1120 CP(1,2) Constitutive parameter at matrix
(F10.0) (psi) position (1,2)
2130 CP(2,2) Constitutive parameter at matrix
(F10.0) (psi) position (2,2)
3140 CP(3,3) Constitutive parameter at matrix
(F10.0) (psi) position (3,3)
4150 THETA Angle of the material axis with
(F10.0) (deg) respect to the global xaxis
Notes
(26)
Card 2D, ITYP = 3, Duncan soil model
Columns
(format)
0105
(15)
0615
Variable
(units)
NON
RATIO
Entry Description
Maximum number of iterations
Default = 5
Moduli averaging ratio
Default =0.5
Notes
(27)
(28)
Go to cards 3D and 4D if MATNAM = USER. Otherwise input is complete for
Duncan model.
160
Card 3D, Hyperbolic parameter for tangent Young's modulus
Columns
Variable
(format)
(units)
0110
C
(F10.0)
(psi)
1120
PHIO
(F10.0)
(radians)
2130
DPHI
(F10.0)
(radians)
3140
ZK
(F10.0)
4150
ZN
(F10.0)
5160
RF
(F10.0)
Entry Description Notes
Cohesion intercept (29)
Initial friction angle
Reduction in friction angle for
a 10fold increase in confining
pressure
Modulus number, K
Modulus exponent, N
Failure ratio, R_
Card 4D, Hyperbolic parameters for tangent bulk modulus, or constant
Poisson ratio option.
Entry Description Notes
Bulk modulus number, K, (30)
Bulk modulus number, M
Poisson 1 s ratio. If a nonzero value
is entered, the bulk modulus is not
used. Instead, the specified constant
VT is used.
Columns
Variable
(format)
(units)
0110
BK
(F10.0)
1120
BM
(F10.0)
2130
VT
(F10.0)
161
i
i
Card 2D, ITYP = 4 (MATNAM = USER), Overburden dependent model,
user defined table, repeat Card 2D as needed to define
input table, last card is blank to terminate reading.
Columns
Variable
(format)
(units)
0110
H(N)
(F10.0)
(psi)
1120
E(N)
(F10.0)
(psi)
2130
GNV(N)
(F10.0)
Entry Description
Overburden pressure for
table entry N
Young's secant modulus for
table entry N
Poisson's ratio table entry N
Notes
(31)
* Note, Card 2D is not required if MATNAT is other than USER since
overburden dependent tables are stored in CANDE for specified
categories of soil.
ITYP =5, and MATNAM = GRAN, MIXE, or COHE; ExtendedHardin
Card 2D.
ITYP = 5, a
model for t
Columns
Variable
(format)
(units)
0110
XNUMIN
(F10.0)
1120
XNUMAX
(F10.0)
2130
XQ
(F10.0)
3140
V0IDR
(F10.0)
4150
SAT
(F10.0)
5160
PI
(F10.0)
Entry Description
Poisson's ratio at low shear strain
Default =0.10
Poisson's ratio at high shear strain
Default =0.49
Shape parameter q for Poisson's ratio
function
Default =0.26
Void ratio of soil, range 0.1 to 3.0
Ratio of saturation, range 0.0 to 1.0
Plasticityindex/ 100, range 0.0 to 1.0
Notes
(32)
6165
NON
Maximum iterations per load step;
Default = 5
162
Card 2D. ITYP ■ 5, and MATNAM = TRIA; ExtendedHardin model for
triaxial data input
Entry Description
Same as card above (XNUMIN, XNUMAX, XQ)
Hardin parameter used to calculate
maximum shear modulus
Hardin parameter used to calculate
reference shear strain
Hardin parameter used to calculate
hyperbolic shear strain
Maximum iterations per load step
Default = 5
Columms
Variable
(format)
(units)
0130
3140
SI
(F10.0)
4150
CI
(F10.0)
5160
A
(F10.0)
6165
NON
(15)
Card 2D. ITYP = 6, interface property definition
Columns
Variabl
(format)
(units)
0110
ANGLE
(F10.0)
(deg)
1120
FC0EF
(F10.0)
2130
TENSIL
(F10.0)
(lb/in)
Entry Description Notes
Angle from xaxis to normal (33)
of interface
Coefficient of friction
Tensile breaking force of contact
nodes
* * * End of input * * *
163
COMMENTARY NOTES
(1) Each problem begins with the command ANALYSIS. The DESIGN option
is not available for box culverts. The program will continue to
execute problem data sets backtoback until the command STOP is
encountered.
(2) Setting LEVEL=2 signals the program that the automatic mesh
generation feature will be used. "Level 2 box" is distinguished
from "Level 2 pipe" by a subsequent instruction in Section B.
Setting LEVEL=3 allows description of arbitrary reinforced concrete
structures and loading conditions. LEVEL=1 is not operable for
box cuJ verts.
(3) By setting PTYPE = CONCRE, the reinforced concrete beamrod element
is used to model the culvert (for other pipe types see 1976 manual).
(4) Setting PDIA = 1.0 signals the program that Section B input data
is for a concrete box as opposed to a concrete pipe. Also, if
LEVEL = 2, it subsequently signals the program to read Section C
input for "level 2 box" instead of "level 2 pipe".
(5) Defining the default concrete wall thickness, PT, is simply for
convenience in limiting input data on subsequent cards.
(6) RSHAPE controls the two options for defining section properties
around the box. Setting RSHAPE = STD allows simplified input
for standard ASTM box sections and can only be used with LEVEL = 2.
Setting RSHAPE = ARBI allows the user to arbitrarily define section
properties at each node around the box and may be used with LEVEL =
2 or 3.
(7) Generally set NONLIN = 3 for all problems. Other options are pri
marily for behavior studies.
(8) Figure B.3 illustrates the concrete material parameters representing
the concrete stressstrain behavior. Generally, the default options
provide reasonable parameter values except for cracking strain
164
STNMAT(l) and compressive strength PFPC which are conservative.
Cracking strain values up to 0.0001 were used for box culverts
studied in this report.
(9) Figure B.4 illustrates the reinforcement material parameters
representing the steel stressstrain behavior. Default options
provide reasonable parameter values except for steel yield stress
which is conservative. Standard ASTM box section reinforcement
assumes 65,000 psi yield strength.
(10) The spacing parameter is used only for crackwidth predictions in
the GergelyLutz formula (see Chapter 3). The default value was
used in this study.
(11) For the RSHAPE = STD option, refer to Figure B.5 for illustration
of standard box section parameters. If the top, side and bottom
slabs are the same thickness, these input variables can be skipped
and the default value PT, input on Card IB, will be used.
(12) Haunch dimensions are used by CANDE to increase the wall thickness
at corner nodes by a simple averaging process and are shown on
the printed output. Generally, HH = HV = PT.
(13) Steel placement is illustrated in Figure B.5 and corresponds to
standard ASTM box designs. All reinforcement steel areas are to
be defined per inch of length in the longitudinal direction. Con
crete cover to all steel centers is specified with the parameters
TC. If variable TC values are desired use RSHAPE = ARBI.
(14) For the RSHAPE = ARBI option, refer to Figure B.6 for illustration
of parameters. In the level 2 option, the section properties are
defined individually at the 15 points (nodes) shown in the figure.
For level 3 solutions, the section properties are defined at the
culvert nodes (NPPT) established by the user.
(15) The embankment and trench configurations are illustrated in
Figure B. 7 and B.8. Each is composed of three soil zones; in situ,
bedding, and fill.
165
(16) By setting W0RD1 = MOD, the level 2 mesh can be selectively
modified by using the extended level 2 option. Modifications
include; defining new soil zones and shapes and specifying live
loads. When this option is exercised, additional data cards C4 to
C7 are inserted after card C2. Card C3 does exist for level 2
box input. Input instructions for cards C4 to C7 are in the 1976
user manual and the finite element mesh topology for level 2 box
is shown in Figure B.ll to B.14.
(17) Construction increments for the trench and embankment installations
are shown in Figures B.9 and B.10. In both cases, the first con
struction increment contains the box culvert and in situ soil.
Increments 2,3,4 are each composed of two rows of elements uni
formly spaced along the sides of the box. Increments 5 to 9 are
composed of one element row, increment 5 is 1/3 R2 thick, and
increments 6 to 9 are 2/3 R2 thick. For deep fill heights, sub
sequent increments are formed with equivalent overburden pressure
(see note 19). The special case of NINC = 1 combines all incre
ments into one (not recommended).
(18) Rl and R2 define the box size and control the overall dimensions
of the mesh as shown in Figure B.ll.
(19) HTCOVR is the distance from the middle of the top slab to the
final soil surface. If HTCOVR is specified greater than 3R2,
the mesh top boundary is truncated at the 3R2 level, and the
remaining soil load is applied as equivalent increments of over
burden pressure (i.e. DENSTY * (HTCOVR  3R2)/(NINC  9)). If
HTCOVR is specified less than 3R2, the horizontal mesh line
closest to HTCOVR is moved to the specified height, but with the
condition that at least two layers of soil exist over the top
of the box.
(20) TRNWID defines the trench width from the middle of the box sidewall
to the in situ soil as shown in Figure B.7. The vertical mesh
line (Figure B.ll) closest to the specified position is moved to
this position to form the trench wall boundary. Minimum value
166
for TRNWID is 0.1 Rl. If TRNWID is greater than 4R1 an embankment
installation is obtained.
(21) BDEPTH defines the depth of bedding below the bottom slab. The
bedding zone is composed of one layer of elements and is kept within
the depth limits (1/10) R2 to (2/3)R2. The bedding width extends
one element beyond the box side.
(22) For level 2 box, three sets of D cards are to be input corresponding
to the predefined soil zones: I = 1, 2, and 3 implying in situ,
bedding, and fill, respectively. For level 3, I corresponds to
material number of element defined by user.
(23) Any soil model (ITYP = 1, 2, 3, 4 or 5) may be assigned to any
soil zone. Choice of a soil model is dependent on the problem
objective, availability of actual soil data, and user's preference.
Suggestions for soil model applications are given in subsequent
notes.
(24) The MATNAM subcategories provide a simplified data input option
for ITYP = 3, 4, and 5 wherein the soil model parameters for
standard types of soil are stored in the CANDE program. Alternatively,
by setting MATNAM = USER (or MATNAM = TRIA for ITYPE = 5) model
parameters may be defined by the user.
(25) The linear elastic model (ITYP = 1) is useful for parameter studies
and bracketing solutions with soft and stiff moduli values. See
Chapter 7 (Table 7.2) for typical moduli values. It is generally
reasonably to model in situ soil with the elastic model.
(26) Orthotropic models can be used to simulate reinforced earth (see
Reference 31).
(27) If the maximum number of iterations (NON) for convergence is ex
ceeded, the program advances to the next load step. If NON is
specified as a negative value, iteration values and convergence
checks are printed out.
(28) Generally set RATIO = 0.5. If Duncan model is used for preexisting
soil zones (e.g. bedding and in situ), set RATIO = 1.0. The Duncan
167
model is probably best suited for characterizing fill soil. See
Chapter 9 (Table 9.2) for hyperbolic parameters corresponding to
standard soil types stored in the program.
(29) For MATNAM = USER, the tangent Young's modulus hyperbolic para
meters are input by the user (usually determined from triaxial
tests, see Reference 29).
(30) For MATNAM = USER, the tangent bulk modulus hyperbolic parameters
may be specified. Or, as an alternative, a constant Poisson's
ratio may be specified. The latter option is the original version
of the Duncan soil model, still preferred by some investigators.
(31) Moduli values for the overburden dependent model correspond to
secant relations from confined compression tests. Thus, this
model provides reasonable representation of soil behavior in
zones where deformation is primarily vertical. For this reason,
the model is better suited for rigid culvert installations than
flexible culvert installations. If MATNAM is other than USER, the
table entries are automatically supplied by CANDE. Table values
are listed in the 1976 user manual, page 39.
(32) The Hardin model is discussed in detail in the 1976 CANDE manual.
This option is best utilized in conjunction with triaxial test
data.
(33) See the CANDE 1976 manual.
168
nitia' crushing
tensile njpture
Flgufie B.3  Idealized StressStrain Diagram of Concrete.
f«
PFSY
>S S
Figure B.4  Idealized StressStrain Diagram of Steel.
169
AS2
HH
2 R2
Hvt
AS3 / t
L 2_£1
LI
\
/—
t
PTT
\, .
 rTC
<PTS
AS1_^"
\_
PTB
/
r
AS4
XL1 =
JJ.
R1
Figure B.5  Box Culverts Parameters for RSHAPE  STD
and Level 2 Solution.
2 R2
2
3
4 5
9 •
(
\
+6
7)
.8
15 14 13 12 11
R1 L
TBI
A
/
Asiy
k
K
aso
'BO
PTV
+
Figure B.6  Box Culvert Generated Mesh (Level 2), and Properties
Definition (Level 2 or 3) .
170
equivalent
overburden
soil pressure
3R2
2R2
3R2
Figure B.7  Trench Soil Installation.
3R2
2R2
/ f equivalent
overburden
^ soil pressure
3R2
Figure B.8  Embankment Soil Installation.
171
3R2
R2
R2
3R2
truncated soil equivalent
jS over burden pressure (n9)
V y V
Figure B.9  Soil Layers and Construction Increment when they
are Applied for Trench Soil Installation.
truncated soil equivalent
^x" over burden pressure (n9)
3R2
R2
R2I
3R2
Figure B.10 Soil Layers and Construction Increment when they
are Applied for Embankment Soil Installation.
172
R2
fR2
•R2
•§R2
R2
R2
R2
R2
§R2
¥ 2
'
^
A
►X
>
I<
^BOEPT, 1 
+
*' jsigkigLLi^
R1
fP',?"'p"'  ™ 1
R1
f
Figure B.ll Geometry of the Soil Undeformed Grid Configuration.
173
ELEMENT NUMBERS
14
14,
i m:
1 14
1 14
5 146
147
143
149
150
13
n;
! 13:
I 13'
1 13
5 13S
137
133
139
140
12
12;
! 12:
1 12'
1 12
5 125
127
123
129
~130
11
ti,
1 n.
i 11
1 11
5 115
117
113
119
120
10
loi 10:
1 10
1 10
5 iOS
107
103
109
no
95
96
97
99
39
100
89
90
91
92
93
94
53
94
9S
35
57
99
77
73
79
90
91
92
71
72
73
74
75
7S
SS
56
57
59
59
70
55
56
57
S3
59
SO
51
52 '
S3
54
US
>45
47
49
49
50
SI
52 *
S3
54
35
35
37
39
33
40
41
42
43
44
25
25
27
23
29
30
31
32
33
34
IS
15
17
19
19
20
21
22
23
24
Figure B.12  Soil Mesh Elements Number.
174
NOCfiL NUM2ESS
157
133
153
150
151
132
153
1S4
153
136
145
u7 .
,14 3
143
ISO
1S1
152
153
iSU
155
1 35
135
137
133
133
140
141
142
143
i44
«24
125
125
127
123
123
130
131
132
133
113
ill
115
115
117
113
113
120
121
122
102 103 1 C4 105
1CS
107
103
103
no
111
35
35
37
S3
33
100
33
39
30
31
32
33
81
32
23
Sm
35
35
71 175
75
77
73
73
57
S3
S3
70
71
72
55
57
53
55
SO
61
52
53 .
54
55
4S
h5
47
43
49
SO
Si
S2
S3
5n
34
35
33
37
33
33
40
41'
42
43
23
24
25
25
27
23
23
30
31
32
12
13
14
15
IS
17
13
13
20
21
1 2 3 4 S 5
10
11
Figure B.13  Soil Mesh Nodal Numbers.
175
INCflEn£NT NUtt££R5
f*
' / i < t
////A
/ / ,; / /
/ /'/ /
t ■ t f *(&
/ /t / '
/ / 1/ /
// / / /
/ / / /
/ y /
k /// /
^77?
/ / / /
f*
y / 'i / /
/ / / /
/
z ' / i ' sv ; / * >
TTi 7
I—77
iXi
/
/
; /
/ / /
/
/ / / , I / /
/
/ /
I /
PT77
/ / /' / /I/ A / / ,/
z 7 ^^ 7
///•'/
7 F7 "~v
/ /l
/ I7TJZZZZZZ7
^77
/
r r
> /
T P~7
/ / / / 1/ / / /
**
/
/i
v 7 7
/*
n
/
/
/ /
' /
y / /
' /V /
/ /
/
/
/
A
/
/ /
/ A
/
/
/
/ L
v'/V'
1 /
t tit
/
/ / /
/ /
■/ / ' / A
? 7 — 7— 7— TrP
V ' ' — L
/ A
V
/
/
/
A
7
/
•/
/
/
/ /
/
7 / T
' /•
/ A
/ / /
1 r
/
d/
/
' /
/
1/ 1 /j
/
/
/
1 <
/
/
/
' / '
/
/
/ V / / / /
< / / / /« /
/
T
/
/ / / / /
/i
/
/ 7
/
/
/:
/J—i
/I
/■/
•m m* rr^ ^r nTTTT tn
' / 1
' /
/ / /
/ /
7 /
/
/ /
' / ./ /
r 7 A
///
/ /
/ /
/ / / / /
/ . / / / /
A
w
A
^
Figure B»1A  Soil Layers Incremental Loading for Trench Installation
and Mesh Boundary Conditions.
17b
APPENDIX C
SAMPLE OF INPUT DATA AND OUTPUT
The three sample problems presented here cover the three solution
levels available when analyzing a reinforced concrete box culvert, that
is: level 2 box, extended level 2, and level 3. The three samples cor
respond to box culverts analyzed during the process of this work.
Table D.l gives a brief description of the box type, solution level,
installation type, and some special comments of the sample problems.
Each problem is presented in the following format: (1) a listing of all
the input cards, and (2) selected CANDE output for the box responses.
The soil responses are not presented.
177
TABLE C.l Example Problems for Analysis of Reinforced
Concrete Box Culverts
Problem
No.
Solution
Level
Soil
Installation
Special Comments
1
2
Embankment
8*68 Box Culvert
ASTM H = 10 ft.
stiff linear soil
soil dead load only
automatic mesh generation
2
2
Extended
Embankment
8*68 Box Culvert
ASTM H = 2 ft.
stiff linear soil
soil plus twice HS20 L.L.
automatic mesh generation
3
3
N
6*42 Box Culvert
outof ground loading
user's input mesh
178
Problem 1  Input
CARD TYPE 1...X...10....X...20....X...30....X...40....X...50....X...60....X...70....X...80
CARD 1A
CARD IB
CARD 2B
CARD 3B1
CARD 3B2
CARD 2C
CARD ID,
CARD 2D
CARD ID
CARD 2D
CARD ID
CARD 2D
150.0
65000.0
B.O
8.0
.01583
0.50
ANALYS 2 CONCRE BOX CULVERT 8*68 (MEDIUM REINFORCED  H=10 FT)
1.0 8.0 STD 3 0.0001
5000.0
8.0 8.0 8.0
.01667 .02417 .02583
EMBA EMBANKMENT  STIFF SQIL
1 3 9 52.00
0.0 INSITUSOIL
0.33
0.0 BEDDINGSOIL
0.33
120.0 FILLSOIL
0.33
STOP
1 1
3333.0
2 1
6666.0
3 1
3333.0
40.00
10.00
120.00
12.00
179
Problem 1  Output
*** PROBLEM NUMBER 1 *«*
BOX CULVERT. 8*68 J MEDIUM REINFORCED  M«10 FT)
EXECUTION MOUE ANAL
SOLUTION LEVEL F.C.AUTD
CULVERT TYPE CONCRETE
♦♦•NEGATIVE PIPE UIAMETEF IMP. IES NEW CANOE OPTION FOR VARIABLE CONCRETE THICKNESS. ***
•♦•OPTION IS RESTRICTED TO ANALYSIS CNLY WITH LEVEL 2BOX, OR LEVEL 3. ***
PIPE PROPERTIES ARE AS FULL3WS ...
(UNITS ARE INCHPOUNO SYSTEM )
NOMINAL PIPE DIAMETER 1.0000
CONCRETE COMPRESSIVE STRENGTH 5000.0000
CONCRETE ELASTIC MODULJS 4266826.00
CONCRETE PUISSON RATIO 0.1700
DENSITY OF PIPE IPCF) 150.0003
STEEL YIELD STRENGTH 65003.0000
STEEL ELASTIC MODULUS 29000000.0
STEEL POISSON RATIO 0.3000
NONLINEAR CUOE (1,2, OR 3 1 3
CONC. CROCKING STRAIN ( 1.2,3) 0.000100
CONC. YIELDING STRAIN (2,3) 0.000566
CONC. CRUSHING STRAIN (2,3) 0.002000
STEEL YIELDING STRAIN (3) 0.002040
SPACING LONGITUDINAL REINFORCEMENT
2.00
180
Problem 1  Output (continued)
NODE *
STEEL
AREAS (IN2)
STCEL
CUVERSUN)
TH1CKNESSUN)
V
ASI(N)
ASO(N)
THI inj
TBOIN)
PTV(N)
1
0.0242
0.0
1.2500
1.2500
8.0303
2
0.02*2
0.0
1.2500
1.2500
3.3000
3
0.0242
0.0167
1.2500
1.2500
8.0000
4
0.0242
0.0167
1.2 500
1.2500
8.3333
5
0.0200
0.0167
1.2500
1.2500
15.0300
6
0.0158
0.0167
1.2 500
1.2500
8.0000
7
n.0150
0.0167
1.2500
1.2500
3.3333
a
0.0158
0.0167
1.2 500
1.2500
8.3000
<>
0.0158
0.0167
1.2 500
1.2500
8.3033
13
0.0158
0.0167
1.2 500
1.2500
8.0000
11
0.0208
0.0167
1.2 500
1.2500
16.0000
12
0.0258
0.0167
1.2 500
1.2500
8.0033
13
O.0258
0.0167
1.2 500
1.2500
8.0000
14
0.0258
0.0
1.2500
1.2500
fl.0000
15
0.0253
0.0
1.2500
1.2500
8.0000
* * BEGIN GENERATION UF CANNED MESH * *
THE DATA TO BE RUN IS ENTITLED
EMBANKMENT  STIFF SOIL
TYPE UF MESH EMBANKMENT
PLOTTING DATA SAVED
PRINT SOIL RESPONSES 1
PRINT CONTROL FOR PREP OUTPUT 3
NUMBER OF CONSTRUCTION INCREMENTS <J
SPAN Or BOX 104.00
HEIGHT OF BOX 80.00
SOIL ABOVE TOP OF BOX (FT) 10.00
*1ESH HEIGHT ABOVE TOP CF BOX IFT) 10.00
SOIL DENSITY ABOVE MESH (PCF) 120.00
IDENTIFICATION OF MATERIAL ZONE WITH MATERIAL NUMBER
MATERIALZONE MATERIAL NO.
INS ITU 1
BEDDING 2
FILL 3
181
Problem 1  Output (continued)
* * BEGIN PREP OF FINITE ELEMENT INPUT * *
THE DATA TO BE RUN IS ENTITLED
EMBANKMENT  STIFF SOIL
NUMBER OF CONSTRUCT UN INCREMENTS 9
PRINT CONTROL FOR PREP OUTPUT 3
INPUT DATA CHECK
PLOT* TAPE GENERATION
ENTIRE FINITE ELEMENT RESULTS UUTPUT 1
THE NUMBER OF NODES IS 167
THE NUMBER OF ELEMENTS IS 150
THE NUMBER UF BOUNCARY CUNDITIUNS IS 200
MATERIAL CHARACTERIZATION F3* SOILS.
PROPERTIES FOR MATERIAL I ******** INSITUSUIL
DENSITY = 0.0
YOUNGS MODULUS" 0.3333E+0*
POISSUNS RATIO= 0.3300C+00
CONFINED MOD.= 0.4938E+0^
LATERAL COEFF.= 0.<t925fc+00
PROPERTIES FOR MATERIAL 2 ******** OEUOINGSOIL
DENSITY = 0.0
YOUNGS MODULUS* 0.6666E+04
POISSONS RAT 10= 0.3300E+00
CUNFINEU MUD.= 0.9877E+04
LATCRAL CUEFF.= O.^925E+0O
PROPERTIES FOR MATERIAL 3 ******** FILLSOIL
DENSITY = 0.12000E+03
•
YOUMGS MODULUS* 0.3333E+0*
POISSONS RATION 0.3300E+00
CONFINED MOD.= 0.4938E+04
LATERAL COEFF.= 0.4925E+00
182
Problem 1  Output (continued)
STRUCTURAL RESPONSE OF CULVERT FOR LOAD INCREMENT
COORDINATES. DISPLACEMENTS AND CRACK DEPTHS ARE IN INCHES
PRESSURES ARE IN LB/IN**2
MOMENTS ARE IN IN.*LB/IN.
THRUST AND SHEAR ARE IN LO/IN.
NPPT
10
11
XCOORD.
YCOURD.
XOISP.
YOISP.
NPRES.
SPRES.
MOMENT
THRUST
SHEAR
CRACK DEPTH
0.0
40.00
0.0
O.36327E+00
0.82420E+01
0.0
0.55599E*04
0.11517E+03
0.0
0.60566E+01
13.00
40.00
0.48848E04
0.36062fcf00
0.85117E+01
0.16894E+00
0.48633E+04
0.11626E+03
0. 10890E+03
0.0
26.00
40.00
0.9081 7E04
0.35366E+00
0.91547E+01
0.37016E+00
0.27284E*04
0.11977E+03
0.22374E+03
0.0
39.00
40.00
0.13419E03
0.34457E+00
0.10687E+02
0.123UE+00
0.948B0E+03
0. 12297E+03
0.35271E*33
0.0
52.00
40.00
0.16387E03
0.33563E+00
0.S0348E+01
0 . 10693E + 02
0.64476E+04
0.36492E+03
0.14262E+03
0.0
52.00
26.67
0.79563E02
0.33548E+00
0.44441E+01
0.45603E+01
0.46375E+04
0.63 646E+03
0.10731E+03
0.0
52. OC
13.33
0.13600E01
0.33524E+00
0.50773E+01
0. 39506 E+Ol
0.36019E«04
0.69321E+03
0.43829E+02
0.0
52.00
0.0
0.15840E01
0.33498E+00
0.62997E+01
0.28920E+01
0.34689E+04
0.73882E+03
0.32017E+02
0.0
52.00
13.33
0.14765E01
0.33470E+00
0.65201E+OI
0.21085F+01
0.44556E+04
0.77216E+03
0.11748E+03
0.0
52.00
26.67
0.91028E02
0.33438E*00
0. 70390E+01
0.84638E+00
0.66014E*04
0.79186E+03
0.20788E«03
0.60218E + 01
52.00
40.00
0. 44758 E03
0.3341 7E+00
0. 15362E + 02
0.33181E+01
0. 10001E+05
0.56033E+03
0. 16154E403
0.0
t2
39.00 0.27012E03 0.14983E + 02 0.24900E+04 0.4804 9E«03
•40.00 ~0.32325E*03 0.5577lt*00 0.31954E*03 0.0
13
26.03 0.25810E03 0.12640E+02 0.24882E+04 0.30094E+03
■40.00 0.31149E+00 0.94244E»00 0.30979E+03 0.0
14
13.00 0.14314E03 0.11427E+02 0.53298E+04 0.14451E+03
■40.00 0.30161E*00 0.53084E+00 0.30021E+03 0.54325E+01
15
0.0 0.0 0.10805E+02 0.624L1E+04 0.0
•40.00 0.29753E+00 3.0 0.29676E+03 0.58268E+01
183
Problem 1  Output (continued)
STRESSES IN CULVERT WALL (PS I ) FCR LOAD INCREMENT
ELLIP. OP
PT
INNER CAGE
OUTER CAGE
CONCRETE
SHEAR
STEEL
STEEL
COMPRESSION
STRESS
1
O.34614E+05
0.0
0. 13187E+04
0.0
2
0.20485E+04
0.0
0 .46337E+03
0. 16134E+02
3
0. 13956E+04
0.13196E+04
0.25882E+03
0.33146E+02
4
0.52378E+03
0.31524E+03
0.9B963E+02
0.52253E*02
5
0. 10500E+04
0.72864E+03
0. 16826E*03
0.96693E«01
6
0.26324E+04
0. 15131E+04
0. 49515E+03
0. 15897E*02
7
0.22193E+04
0.10012E+04
0.40881E+03
0.6493lE*0l
9
0.21999E+04
0.90183E+03
0.40239E+03
0.47433E+01
9
0.26 704E+04
0.13131E+04
0 .49531E+03
0.17405E+02
10
0.26859E+04
0.41673E+05
0. 17685E*r»4
0.30797E+02
11
0.16246E+ K
0. 11323E+04
0.26039E+03
0. 10952E+02
12
0. 13679C+04
0.82989E +03
0.25867E+03
0.71184E+02
13
0.82146E+03
0. 13772E+04
0.25998E+03
0.44584E*02
I*
0.23769E+05
0.0
0. 13087E+04
0.21408E+02
15
0.3234 8E+05
0.0
0. 14766E* 04
0.0
STRAINS IN THE INNER ANU OUTER FIBER OF THE CULVERT WALL
(ONLY STRAINS FOR COMPRESSION ZLNES HAVE PHYSICAL MEANING)
NPPT
INNER STRAIN
OUTER STRAIN
I
0.13426E02
0.29872 E03
2
0.95622E04
0. 10497E03
3
1. 5 1 603 E 14
0.58631E04
4
0.22419E34
0.15875EC4
5
0.3811 7E04
0.28032E04
6
0. 11217E03
0.77045E04
7
0.92608E04
0.54384E04
8
0.9H53E04
0.5042 JE04
9
0.1122 0E03
0.69614E0't
10
0.'»0063ED3
0.16240E02
11
0.58987E04
0.43540E04.
12
0.58597E04
0.41715E04
13
0.41457E04
0.58895E04
14
0.93886ED3
0.29646E03
15
0.12650E02
0.33450E03
184
Problem 1  Output (continued)
CALCULATED SAFETY FACTORS FUR LOAD INCREMENT 9
STEEL YIELD STRESS / MAX. STEEL STRESS 1.560
CONCRETE STRENGTH / MAX. COMPRESSIVE STRESS .... 2.827
WALL SHEAR CAPACITY / MAX. SHEAR 1.987
PERFORMANCE FACTORS
0.C1 INCH / MAX. CRACK WIDTH 1.215
+ * * * NORMAL EXIT FRUM CANDE * * * *
185
Problem 2  Input
CARD
TYPE
1
CARD
1A
ANALYS 2 CONCRE BOX CULVERT 3*63
 HS20
LIVE LOAD
(MINIMUM
CARD
IB
1.0
8.0 STD 3
0.0001
CARD
223
5000.0
150.0
6500
CARD
3B1
8.0
8.0 8.0
8.0
8.0
CARD
332
.02583
.03833 .02917
.01583
0.65
CARD
1C
EMBA EMBANKMENT  STIFF SOIL
CARD
2C
1
3 8 52.00
40.00
2.00
120.00
CARD
4C
2
CARD
7C
124
60
.0
7
CARD
7C
124
51
.1
8
CARD
ID
1 1
0.0 INSITUSOIL
CARD
2D
3333.0
0.33
CARD
ID
2 1
0.0 BEDDINGSOIL
CARD
2D
6666.0
0.33
CARD
ID
L
3 1
120.0 FILLSOIL
CARD
2D
3333.0
0.33
.70....X...80
MOD
12.00
STOP
186
,
Problem 2  Output
• •• PROBLEM NUMBER 1 **♦
BOX CULVERT 8»68  HS20 LIVE LOAD (MINIMUM SUIL CUVERI
EXECUTION MODE ANAL
SOLUTION LEVEL F.E.AUTO
CULVERT TYPE CUNCRETE
••♦NEGATIVE PIPE OIAMETER IMPLIES NCW CANOE OPTION FOR VARIABLE CONCRETE THICKNESS. *•*
•♦•OPTION IS RESTRICTED TO ANALYSIS CNLY WITH LEVEL 2BUX, OR LEVEL 3. ***
PIPE PROPERTIES ARE AS FOLLOWS ...
(UNITS ARE INCHPOUND SYSTEM »
NOMINAL PIPE OIAMETER I. OOOO
CONCRETE COMPRESSIVE STRENGTH 5000.0000
CONCRETE ELASTIC MUDULJS 4286826. 00
CONCRETE PUISSUN RATIU 0. WOO
DENSITY OF PIPE (PCF) 150.0000
STEEL YIELO STRENGTH 65000.0000
STEEL ELASTIC MODULUS 29000000.0
STEEL POISSUN RATIO 0.3000
NONLINEAR COOE (1,2, OR 3) 3
CONC. CRACKING STRAIN (1,2,3) 0.000100
CONC. YIELDING STRAIN (2,31 0.000566
CONC. CRUSHING STRAIN 12,3) 0.002000
STEEL YIELOING STRAIN (3) 0.002040
SPACING LONGITUDINAL RE I NFCPCEMENT
2.00
187
Problem 2  Output (continued)
13DE *
STEEL
AREASUN2)
STEEL
COVERSUNI
THICKNESS!
N
ASIIN)
ASO(N)
TL> I ( N )
TMMN.I
PTV(N)
1
0.0383
0.0
1.2500
1.2500
8.0000
2
0.0383
0.0
1.2500
1.2500
8.00 33
3
0.03 A3
0.0258
1.2 500
1.2503
8.0333
4
0.0333
0.0258
1.2 500
1.2500
8.0303
5
0.0271
0.0258
1.2500
1.2500
1ft. 0000
6
0.0158
0.0258
1.2500
1.2500
9.3303
7
0.0158
0.0258
1.2 500
1.2 500
8.0C00
8
0.0158
0.0258
1.2500
1.2500
8.0000
9
0.0158
0.0258
1.2500
1.2 500
8.0000
10
0.0158
0.0258
1.2 500
1.2500
fl.OCOO
11
0.0225
0.0258
t.2 500
1.2500
16.0033
12
0.0292
0.0258
1.2500
1.2500
8.0003
13
0.0292
0.3258
1.2 500
1.2 500
8.0000
14
0.0292
0.0
1.2500
1.2503
8.0333
15
0.0292
0.0
1.2500
1.2500
8.0000
• * BEGIN GENERATION OF CANNEC MESH * *
THE DATA TO BE RUN IS ENTITLED
EMBANKMENT  STIFF SOIL
TYPE 3F MESH EMBANKMENT
PLOTT ING DATA SAVED
PRINT SOIL RESPONSES I
PRINT CONTROL FOR PREP OUTPUT 3
NUMBER OF CONSTRUCTION INCREMENTS 8
SPAN OF BOX 134.00
HEIGHT OF BOX 80.00
SOIL ABOVE TOP OF BOX (FT) 2.(0
^ESH HEIGHT ABOVE TOP CF BOX (FT) 2.00
SOIL DENSITY ABOVE MESH IPCF1 120.00
IDENTIFICATION OF MATERIAL ZONE WITH MATERIAL NUMBER
MATERIALZONE MATERIAL NO.
INS ITU 1
BEDDING 2
FILL 3
188
Problem 2  Output (continued)
* * BEGIN PREP OF FINITE ELEMENT INPUT * *
THE DATA TO BE RUN IS ENTITLED
EMBANKMENT  STIFF SOIL
NUMBER OF CONSTRUCTION INCREMENTS
PRIMT CONTRCL FOR PPEP OUTPUT 3
INPUT DATA CHECK
PLOT TAPE GENERATION
ENTIRE FINITE ELEMENT RESULTS OUTPUT I
THE NJMEER OF NOOES 13 13*
THE NUMBER OF ELEMENTS IS 120
THE MUMBER OF BOUNCARY CONDITIONS IS 200
* * * CHANGES TO STANCARO LcVEL 2 MfcSH * * * *
* NUMBER OF NOOES TO BE CHANGED 0*
* NUMBER OF ELEMENTS TU BE CHANGED 0*
* ADDITIONAL BOUNOARY CGNUITIONS 2*
******************* **•*
♦ ♦•ADDITIONAL BOJNCARY COND IT IONS . . .P 3 R CES = LHS t DISPLACEMENTS = INCHES..
BOUNOARY
LOAD
XFORCE 0*
YFORCE OR
XY PU TAT I UN
NOOE
STEP
XOISPLACEMENI
YDISPLACEMENT
UEGFEES
12*
7
F a 0.0
F = 0.6000C*02
).">
12*
8
F » 0.0
F = 0.5U0E+02
0.0
189
Problem 2  Output (continued)
MATERIAL CHARACTER IZAT 1UN FOR SUILS.
PROPERTIES FOR MATERIAL 1 ♦*«»**♦* INSITUSLUL
DENSITY = 0.0
YOUNGS MO0ULUS= 0.3333EKK
POISSONS RAT 10= 0.3300C*00
CONFINED MOl).= 0.4938E*04
LATERAL COEFF.= U.4925E+00
PROPERTIES FOR MATERIAL 2 *♦*♦*»♦♦ BELUINGS01L
DENSITY = 0.0
YOUNCS MODULUS" 0.6666E*04
POISSONS RATIO= 0.3300E*00
CONF INEO MOO." 0.9877E+04
LATERAL COEFF.= 0.49251*00
PROPERTIES FOR MATERIAL 3 *♦*••*♦♦ TILLSOIL
DENSITY = 0. 12U00E*33
YOUNGS MODULUS" 0.3333CKK
POISSONS PATIO= 0.3300E*00
CONF INED MOO." 0.4938E*04
LATERAL COEFF." 0.<t925l*00
190
Problem 2  Output (continued)
STRUCTURAL RESPONSE Of CULVERT FOR LOAD INCREMENT 8
COORDINATES. 01 SPLACEMENTS AND CRACK DEPTHS ARE IN INCHES
PRESSURES ARE IN LB/IN**2
MOMENTS ARE IN IN.*LB/IN.
THRUST ANO SHEAR ARE IN LD/IN.
NPPT XCOORO. XOISP. NPRES. MUMENT SHEAR
YCOORO. YDISP. SPRES. THRUST CRACK DEPTH
1 0.0 0.0 0.82803C*0l J. 40991 fc *0'. J. 3
40.00 0.1930<JE*00 0.0 0.25164E*02 0.0
2 13.00 0.89571EG5 0 . I 5124E* 01 0.33993E+04 0.96151C*02
40.00 0.19143E*00 0.17414E*01 0. 13845E*02 0.0
13
11
13
14
15
26.00
0.98505CG5
0.34548E*0l
0. 159«HE *04
0. 1609'»E*03
40.00
0. 18703E*00
3.1<>975E*01
0. 104?JE*02
0.3
3<».00
0.16997E05
0.23285E*0l
0.78224E+03
0.19853E + 03
40.00
0.1«138E*00
3.42829E*r»0
3.26?25C*32
3.0
52.00
0.51747E05
0. 16996E*01
0.35617EO4
0.85171E*32
40.00
0.17587E*00
3 ,33O03E*Ol
0.144 70L : *03
3.0
52.00
0.50454C02
0.9 7901C*00
0.29°05E*0't
O.36796E+02
26.67
0.17581 fc*00
0.20297E*01
0.27393E+03
0.3
52.00
0.84971E02
0. 113 26E*01
0.258711*04
0.22713C*02
13.33
0. 17571E+00
3.23598C+01
0.30319E+02
0.0
52.00
0.95969E02
0.22470E*Ol
0.2 3*4OE*04
0. 18745E*C0
0.0
0.1 7559E*CJ
0.20561O01
3.33263L03
0. J
52.00
0.84966E02
0.23406C* r >l
0.25H21E H4
3.3 )?V/F*J?
13.33
0. 17546E*00
0.2056^C*Ol
0.36005EO3
0.0
52.50
0.531 64CC2
3. 317131*11
o. »i95it*04
O.67l4in»07
26.67
0.17533E*00
0. 155541: ♦Ol
0.3 0413C*03
0.0
52.00
0.16512C03
0.792 491* 11
0.<.3 75">C»0'»
■1.101250*03
40.00
O.I /523E*00
3 .2275'.C»ni
0.26?fc9t*93
0.0
26.00
0.079 13E04
0.627I6E*Ol
3.19106E*04
3.15253E+03
40.00
0.I6368E*00
0.^.0764E*no
0.12 H9L'*03
0.0
13.00
0.43730E04
0.57943E+O1
0.33603E+04
0.74104E*02
•40.00
0.15934E+00
0.20350E*00
0.123Z2E*03
0.0 ,
O'.O
0.0
0.56062E+01
0.383Z2C*04
0.3
•40.00
0.15774E*00
3.0
0.12190E*03
3.3
191
Problem 2  Output (continued)
STRESSES IN CULVERT WALL (PS I ) FOR LUAO INCREMENT
NPPT
1
2
3
'♦
5
6
7
8
9
10
11
12
13
1*
15
ELL1P. OR
INNER CAGE
STEEL
0.17e47E*04
0. I 474 1E*04
3.67317E03
0.35505E*03
C.54772E+03
0.15723E + 04
0. 14181E*04
0.1353 7E04
0. 14655E+C4
0.17599E+04
3.71466E*03
0.37275E+03
0.72027E*03
0. 13637E04
0.15715E*04
OUTER CAGE
STLbL
0.0
0.0
O.7O528C*03
0.31755E*03
0.42065E »03
0. 10729C+04
0.870BOE+03
0.75670E + 03
0.81929E*03
0.10670E+04
0.47993E«03
0.1482 3E*03
0.943 74£*03
0.0
0.0
CONCRETE
;OMPRESSION
■0. 37247E+03
•0. 30974E + 03
■0. 14109E*03
■0.70357E* 02
3.8829lE*02
0.30107F>03
0.26849E*03
0.2 53 < ;5E»O3
O .27494E*03
0.33277E* 03
3. U432E*0j
0.68036E*02
3. 1831 IF*03
0.32448E*OJ
0 .36778E* 03
SMEAR
STRESS
0.0
0. 14245E+02
0.23843E*02
0.2 Wl2fc*02
0.57743E*0I
0.54512E«01
0.3365fcb»01
J.27771G01
0.450321*01
0.99471F.»0l
0.68642E+01
0. 35853L«02
0.22597C»02
0.1 197BC»02
0.0
STRAINS IN THE INNER AND OUTER FIBER OF THE CULVERT WALL
(ONLY STRAINS FOR COMPRESSION ZCNES HAVE PHYSICAL MEANING)
NPPT
INNER STRAIN
OUTER STRAIN
I
0.8 t9<J9E0 4
0.843 75E
■C4
2
0.6781 5E34
0.70166E
■04
3
0. 30954E04
0.31962E
■04
4
0.15938E04
0.14761T
■C4
5
0.20001E34
0. 16013E
■04
6
3.68232E34
0.52530E
■04
7
0.60822E04
0.43649C
■04
8
0.57529E04
0.38795E
04
9
O.A2283E04
0.420U4E
■04
10
0.75384E04
0.53641C
C4
11
0.25897E14
3. 10531b
■04
12
0. 15412E04
0.83669E
C5
13
0.34469E04
0.41481 E
■04
14
3.64327E34
0. 73535 E
•04
15
0.73873E04
0.83314E
■04
192
Problem 2  Output (continued)
CALCULATED SAFETY FACTORS FOR LOAD INCREMtNT ft
STEEL YIELO STRESS / MAX. STEEL STRESS 36.*21
CCNCRETE STRENGTH / MAX. COMPRESSIVE STFESS .... 13. *24
WALL SHEAR CAPACITY / MAX. SHEAR 3.9**
PERFORMANCE FACTORS
0.01 INCH / MAX. CRACK WIDTH 9909.996
* • * * NORMAL EXIT FROM CA>iDE * • * *
193
Problem 3  Input
CARD
TYPE
CARD
1A
ANALYS
3 CONCRE BOX CULVERT
TEST
(6*42)
TEST1
CARD
IB
1.0
7
.0
ARBZ
3
0.0001
CARD
2B
6965.0
.17
150.0
99430.0
CARD
3B
.03475
.00000
1.'
443
7.375
CARD
3B
.03475
.03550
1.'
443
7.375
CARD
3B
.03475
.03550
1.
443
7.375
CARD
3B
.03475
.03550
1.443
7.375
CARD
3B
.01737
.03550
14.0
CARD
3B
.00000
.035
50
CARD
3B
.00000
.03550
CARD
3B
.00000
.035
50
CARD
3B
.00000
.03550
CARD
3B
.00000
.035
50
CARD
3B
.01337
.03550
14.0
CARD
3B
.o:
2675
.035
50
1.006
7.438
CARD
3B
.02675
.035
50
1.006
7.438
CARD
3B
.0.
2675
.03550
1.006
7.438
CARD
3B
.0.
2675
.00000
1.006
7.438
CARD
1C
PREP
BOX CULVERT
OUTOFGROUND STUDY
(TEST 6*
CARD
2C
10
3
1
17
IS 14
CARD
3C
1
0.0
27.5
CARD
3C
9.0
27.5
CARD
3C
3
19.0
27.5
CARD
3C
4.
29.0
27.5
CARD
3C
5
39.5
27.5
CARD
3C
6
a
39.5
17.0
CARD
3C
7
39.5
8.5
CARD
3C
8
39.5
0.0
CARD
3C
9
39.5
8.5
CARD
3C
10
39.5
17.0
CARD
3C
11
39.5
27.5
CARD
3C
12
29.0
27.5
CARD
3C
13
19.0
27.5
CARD
3C
14
9.0
27.5
CARD
3C
15
0.0
27.5
CARD
3C
16
6.0
33.5
CARD
3C
L 17
12.0
33.5
CARD
4C
1
2
1
1
CARD
4C
2
3
2
2
CARD
4C
3
4
3
. o
3
CARD
4C
4
5
4
4
CARD
4C
5
6
5
5
CARD
4C
6
7
6
6
CARD
4C
7
8
7
7
CARD
4C
8
.9
8
8
CARD
4C
9
10
9
9
CARD
4C
10
11
10
10
CARD
4C
11
12
11
11
CARD
4C
12
13
12 •
12
CARD
4C
13
14
13
13
CARD
4C
14
15
14
14
■
CARD
4C
L IS
16
17
14
1
.60....X...70....X...80
1415
194
Problem 3  Input (continued)
CARD TYPE 1.. .X...10....X...20....X...30....X...40. ...X...50....X...60.. ..X...70....X...80
CARD 5C 11 0.0 1
CARD SC IS 1 0.0 1
CARD 5C 16 1 0.0 1
CARD 5C 17 1 0.0 1
CARD SC 2 100.0 1
CARD SC 2 100.0 2
CARD SC 2 100.0 3
0.0
0.0
1
0.0
1
0.0
100.0
100.0
100.0
100.0
100.0
100.0
100.0
100.0
100.0
100.0
CARD SC 2 100.0 4
CARD SC 2 100.0 3
CARD SC 2 100.0 6
CARD 5C 2 100.0 7
CARD SC 2 100.0 8
CARD SC 2 100.0 9
CARD SC L 2 100.0 10
CARD ID L 1 1 '
CARD 2D 1.0E+11 0.0
STOP
195
Problem 3  Output
♦ ♦• PROBLEM NUMBER i *•*
BOX CULVERT TEST «6**2J TESTl
EXECUTION MUOE ANAL
SOLUTION LEVEL F.E.USE*
CULVERT TYPE CONCRETE
♦♦♦NEGATIVE PIPE DIAMETER IMPLIES NEW CANOC OPTION FOR VARIABLE CONCRETE THlCKNESi. ♦ ♦*
♦•♦OPTION IS RESTRICTED TO ANALYSIS CNLV WITH LEVEL 2BOX, OR LEVCL 3. ♦♦♦
PIPE PROPERTIES ARE AS FOLLOWS ...
(UNITS ARE INCHPOUND SYSTEM ,)
NOMINAL PIPE DIAMETER I. OOOO
CONCRETE COMPRESSIVE SMENGTh 6965. OCOO
CONCRETE ELASTIC MUOULUS 505*5*4.00
CONCRETE POISSON RATIO 0.1700
DENSITY OF PIPE (PCF) 150.0000
STEEL YIELD STRENGTH 99A30. 0000
STEEL ELASTIC MODULUS ^ 29000000.0
STEEL PCiSSUN RATIO 0.3000
NCNLINEAR CODE (1,2, OR 3) 3
CONC. CRACKING STRAIN (1,2.3) 0.000130
CONC. YIELDING STRAIN (2,31 0.000668
CONC. CRUSHING STRAIN (2,3) 0.002000
STFEL YIELOING STRAIN (3) 0.003120
SPACING. LONGITUDINAL RE I NFORCEMENT
2.00
196
Problem 3  Output (continued)
njoe *
STEEL AREASIIN2)
STEEL COVERS! IN)
THICKNESS! IN »
N
ASIfNI
ASO(N)
TBI IN)
TDCXN)
PTV(N)
I
0.0347
0.0
1.4430
1.2500
7.3750
2
0.0347
0.0355
1.4430
1.2500
7.3750
3
0.0347
0.0355
1.4430
1.2500
7.3750
4
0.0347
0.0355
1.44J0
1.2500
7.3750
5
0.0174
0.0355
1.2 500
I. 2503
14. JOJO
6
0.0
0.0355
1.2 500
1.2500
7. 1003
7
0.0
0.0355
1.2 500
1.2500
7.0000
9
0.0
0.0355
1.2500
1.2500
7.0000
n
0.0
0.0355
1.2500
1.2500
7.0000
10
0.0
0.0355
1.2500
1.2500
7.0000
11
0.0134
0.0355
1.2500
1.2500
14.0003
12
0.0267
0.0355
I .0060
1.2500
7.4380
13
0.0267
0.0355
1.0060
1.2500
7.4380
14
0.3267
0.0355
1.0 ObO
1.2500
7.4380
15
0.0267
0.0
1.0060
1.2500
7.4380
* * 8EGIN PREP OF FINITE ELEMENT INPUT * *
THE OATA TO. OE RUN IS ENTITLED
BOX CULVERT CUTUFGRUUNU STUOY I TEST 6*42)
NUMBER OF CONSTRUCTION INCREMENTS 10
PRINT CONTROL FOR PREP OUTPUT 3
INPUT OATA CHECK
PLUT TAPE GENEFATION
ENTIRE FINITE ELEMENT RESULTS OUTPUT —  1
THE NUMBER OF NOOES IS 17
THE NUMBER OF ELEMENTS IS 15
THE NUMBER OF BOUNCARY CUNDITIUNS IS 14
197
Problem 3  Output (continued)
♦ ••BOUNDARY CONDI T IONS .. .FORCES * LBS DISPLACEMENT S » INCHES...
8OUN0ARY
NODE
1
15
16
17
2
2
2
2
2
2
2
2
2
2
LOAO
XFORCE
0*
Y
FORCE OR
STEP
XOISPLACEMEMT
Y
OlSPLACtMENT
I
D » 0.0
F
3
0.0
1
» 0.0
F
S
0.0
1
F « 0.0
■S
0.0
1
F • 0.0
=
0.0
I
F » 0.0
F
=
O.lOuOE«03
2
F » 0.0
F
a
0.1000L+03
3
F * 0.0
F
s
0.1000C+03
4
F » 0.0
F
a
O.IOOOE*03
5
F » 0.0
F
■
0.1000C*03
6
F = 0.0
F
=
O.IOOOE*03
7
F ■ 0.0
F
a
0.1000E+03
8
F » 0.0
F
=
0.1000C*03
9
F a 0.0
F
3
0.1000E+03
10
F * J.O
F
3
0.1000E*03
XY ROTATION
OE^EES
0.0
0.)
0.0
0.0
O.J
0.0
0.0
0.0
0.0
0.5
0.0
0.1
0.1
O.rt
198
Problem 3  Output (continued)
STRUCTURAL RESPONSE UF CULVERT FOR LOAO INCREMENT 5
COORDINATES t DISPLACEMENTS AND CRACK DEPTHS ARE IN INCHES
PRESSURES ARE IN LB/IN**2
MOMENTS ARE IN IN.*LB/IN.
TH<US7 AND SHEAR ARE IN LB/IN.
NPPT XCOORD. XOISP. NPRFS. MOMENT SHEAR
YCOORO. rDISP. SPRES. THRUST C«ACK DEPTH
1 0.0 0.0 0.6441CE*<>0 0.91441C*04 0.0
27.50 0.15773E*00 3.0 0.l3239E«02 0. 56336E«0l
2 9.00 0.21162E05 3.«3278E»02 0.9ll7flE«0* 0.25597E*03
27.50 3.l4942t*00 3.18370E05 3.13239E*02 3.57301EOI
3 19.00 0.50681E05 0.65051F*00 0.4027OEKK 0.51229EO3
27.50 0.12B14E*00 0.17166E05 0.13239E*02 0.0
4 29.00 0.83627E05 3.62088C*00  J. 1 1200E*0« 0.51872E*03
27.50 0.10406E+00 0.26052C05 0.13239E*02 0.0
5 39.50 0.10792EG4 0.46690t*00 0.66008E*04 3.2675<>C»03
27.50 0.78849E01 0 .461 1 8E*00 0.25778E*03 0.0
6 39.50 0.24313E01 0.34437E02 0.67'»93E*0<» 0.12265E*02
17.00 0.78740E01 3.63980E»03 0.53184E*03 0.47313COI
7 39.50 0.39382E01 0.29297E02 0.6B619E»04 0. 1 32 36C 02
8.50 0.785*6E0l 0.64936E»no 0.53763EO3 0. 474J5E/01
8 39.50 0.45814E01 3.51970O03 3.69743E*r>'. 0.13221FO2
0.0 0.78341E01 0.6452?E*00 0. 5'.3 1 4E *03 0.47867CO1
9 39.50 0.4200JE01 0.37807102 0. 7006 7 E »0<. 0.13231T»02
8.50 0.7bl09E01 0 . 6<.568E»00 0. 5<.P62t »02 0. «. 7<)<,2C ♦ 01
10 39.50 0.26661E01 0.7863VEO2 0. 71 90 9L *04 0.1322'»E*02
17.00 0.77876fc0l 0 .64 LOH^ J 0 . 55'.'. I E 03 J.'.OOl '»C *0 1
11 39.50 0.9O*92LO5 . 45"06F *00 0.73 3801*0'. 0. ? 75 sor »03
27.50 0.77757C01 3.45692t*0J 0. 205 3 5C»03 n.o
12 29.00 0.A619JE05 0.6M27r«O3  3 . 1 '. I 3 I L ♦ V»  3. r ><W S'.OO «
27.50 0.5006HE01 3.ni67l<^. 0. 1 3246C »)2 J.O
11 19.00 0.33335C05 0.64524E*00 0.<»2963E *0<» 0. 57<.06CO3
27.50 0.23575601 0.39101E05 0. 132*<SE»02 0.0
14 9.00 0.57829E06 0.61072E+02 0.10070E*05 0. 287 liC*03
27.50 0.11727E07 3.24769E02 0. 13234E02 . 586 06E Ol
15 0.0 0.0 0.64529E*00 0.100'»3E*05 3.0
27.50 0.92615E02 0.0 0. 13222C*02 0.505?U*Ol
199
Problem 3  Output (continued)
STRESSES IN CULVERT HALL (PS I) FOR LUAD INCREMENT
NPPT
I
2
3
4
5
6
7
8
9
10
11
12
13
14
1 5
ELLIP. OR
INNER CAGE
STEEL
0.47231E+05
0.4832 0E*05
0. 15645E*04
0.42373E*03
3. 10902E+04
0.0
0.0
0.0
0.0
0.0
0.12164E+04
0.65296E+03
0. 19400EO4
0.62375EO5
0.62147E*05
OUTER CAGE
STEEL
0.0
■0.97052E*03
0.16679E*04
0.47940E*03
0.85525E*03
0.27082E«05
0.27558E*05
0.29503E*05
0.29974E+05
0.30441E*05
0.95255E »03
0.56926E+03
0.1 7701 E*04
0.80526E«03
3.0
CONCRETE
COMPRESSION
■3. 247 11C + 04
3.2 31 01 E* 34
0.41377E*03
0. H470f>03
0.21280E+03
0.20249E*04
0.20582EO'.
0.21715E+0*
0.22046E*n<,
0.22376EO4
■0.23742E+03
0. 14554E+03
■3.43570E*O3
0.26233E*04
0.26309E*04
SMEAR
STRESS
0.0
0.42460E+02
0.84 )73C*02
0.86045E*02
0.209C8E*02
0.2306 9E*0l
0.23019L : *01
0.2299'»E+01
3.22<'62E*01
3.22999E«0t
0.21608E*02
0. 89944E*02
■0.90976E*02
0.45513E*02
0.0
STRAINS IN THE INNER AND OUTER FIBER Of THE CULVERT WALL
(ONLY STRAINS FOR COMPRESSION ZONES HAVE PHYSICAL MtANING)
NPPT
INNER STRAIN
OUTER STRAIN
1
0.1 <580E02
0.47430E
03
2
0. 19930E02
3.44340E
03
3
0.80352E04
0.79416E
04
4
0.22031E04
0.22609E
04
5
0.40844E04
0.334 73E
C4
6
0.38864E03
0.11191E
02
7
0.39504E03
0.1138 6E
02
8
0.41679E03
0. 12176L
02
9
0.42314E03
0.12370E
02
10
0.42947E03
3.12563E
02
11
0.45569E04
0.3 72B8L
04
12
0.27935E04
0.27114E
04
13
0.83477E34
0.83626E
04
14
0.23422E02
0.50350E
03
15
0.23341E02
0.50496E
03
200
Problem 3  Output (continued)
CALCULATED SAFETY FACTORS FOR LOAD INCREMENT 5
STEEL YIELD STRESS / MAX. STEEL STRCSS 1.59'*
CONCRETE STRENGTH / MAX. COMPRESSIVE STRESS .... 2.647
HALL SHEAR CAPACITY / MAX. SHEAR 1 . 835
PERFORMANCE FACTORS
0.01 INCH / MAX. CRACK WIDTH 0.897
201
APPENDIX D
CANDE PROGRAM OVERLAY
This appendix provides job control language (JCL) for IBM computers
in order to reduce core storage requirements (region size) for executing
CANDE1980. Instructions and examples are given for two FORTRAN IV com
pilers commonly supported at. IBM installations; the G and the Hextended
(HX) compilers.
For reference, a tree chart of the CANDE1980 subroutines is shown
in Table D.l indicating the calling sequence of all subroutines. Sub
routines CANBOX and DUNCAN are new subroutines added to the CANDE1980
program, and subroutines CONMAT, CONCRE, and READM have been extensively
modified from the CANDE1976 program. Some minor changes have been made
to other subroutines.
Table D.2 gives the JCL to compile an overlayed version of CANDE on
the G compiler presuming the source program resides on a disk file created
by standard TSO operations. Here, the overlay commands (ENTRY MAIN through
INSERT BURNS) provide a simple overlay structure that may be used as a
guide for overlaying CANDE on most computers.
In a similar manner, TABLE D.3 provides JCL for compiling an over
layed version of CANDE with the HX compiler. Here, the overlay structure
is slightly different and takes advantage of special overlay options
(i.e. OVERLAY C (REGION)) available at IBM installations. The overlay
commands in Tables D.2 and D.3 may be interchanged, however the overlay
structure in Table D.3 is more efficient on IBM.
Once a load module is created from either Table D.2 or D.3, the
JCL to execute the program is shown in Table D.4. Efficiency comparisons
of executing a typical problem (Example 1, Appendix C) are shown in
Table D.4 for the G and HX compilers with and without overlay. These
202
examples were executed on the IBM 370/168 computer at the University of
Notre Dame. It is observed that HX compiler with overlay provided the
most efficient results in terms of core storage, execution time, and
total cost.
203
TABLE D.l  Subroutine Tree Structure for CANDE
START
CALLS
CALLS
CALLS
CALLS
CALLS
MAIN
ALUMIN
BASIC
BURNS
CONCRE
EMOD
HINGE
**PRHERO
**PRHERO
CONMAT
INVER
SETU
**PRHERO
PLASTI
**PRHERO
**PRHERO
STEEL
EMOD
HINGE
**PRHERO
INVER
SETU
**PRHERO IS:
PRHERO
ESTAB
HEROIC
PREP
RESOUT
BAKSUB
CONVT
DUNCAN
HARDIN
INTPl
READM
REDUCI
RESPIP
STIFNS
STRESS
CANBOX
CANI
GENEL
GENEND
GENNOD
MODMSH
RESOUT
SAVED
PRINC
CONVT
CONVT
ANISP
CONVT
BEAMEL
STFSUB
XFACES
BEAMST
GEOM
XCAN2
XCAN
AF
SAVEG
PRINC
GEOM
J
204
TABLE D.2  JCL to Create an Overlayed Version
of CANDE Using the G Compiler
//CNDEOVLY JOB (XX,XXXX, , 15) , IDNUMBER,NOTIFY=TSOID#,
// REGION=256K,TIME=l
//STEP1 EXEC FORTGCL,PARM.LKED='OVLY,MAP,XREF,LIST'
//FORT.SYSLIN DD UNIT=DI SK,DISP=(NEW,P ASS ) ,SPACE=(TRK, (10,5) ) ,
// DSN=&&LOADSET,DCB=BLKSIZE=80
//FORT.SYSIN DD UNIT=DISK,VOL=SER=XXXXXX,DSN=TSOID#. CANDE. FORT,
// DISP=SHR
//LKED.SYSLMOD DD SPACE= (1024, (600,50,1) ,RLSE) ,DISP= (NEW, CATLG) ,
// UNIT=DISK,VOL=SER=XXXXXX,DSN=TSOID#. CANDE. LOAD(CANDE)
//LKED.SYSLIN DD UNIT=DISK,DSN=&&LOADSET,DISP= (OLD, DELETE)
//LKED.SYSIN DD *
ENTRY MAIN
INSERT MAIN, PRHERO,RESOUT,PRINC,ESTAB
INSERT STEEL, ALUMIN,EMOD
INSERT HINGE, SETU,INVER
INSERT C0NCRE,C0NMAT
INSERT PLASTI, BASIC
OVERLAY A
INSERT PREP
OVERLAY B
INSERT CANB0X,CAN1,XCAN,XCAN2
OVERLAY B
INSERT GENNOD , SAVEG , GENEL , GENEND , AF , MODMSH , SAVED
OVERLAY A
INSERT HEROIC ,CONVT , GEOM, RESPIP , REDUCI , BAKSUB ,XFACES
OVERLAY B
INSERT READM,ANISP, DUNCAN, HARD IN, INTP1
INSERT STIFNS , BEAMEL , STFSUB , STRESS , BEAMST
OVERLAY A
INSERT BURNS
/*
Note: Overlay commands (ENTRY MAIN through INSERT BURNS)
start in column 2.
205
J
TABLE D.3  JCL to Create an Overlayed Version of CANDE
Using the H Extended Compiler
//CNDEOVLY JOB (XX,XXXX, ,15) ,IDNUMBER,NOTIFY=TSOID#,REGION=256K,
// TIME=2
//STEP1 EXEC F0RTXCL,PARM.F0RT='0PT(2)%PARM.LKED='0VLY,MAP,XREF,LIST'
//FORT.SYSIN DD UNITDISK, VOLSERXXXXXX,DSNTSOID#. CANDE. FORT,
// DISP=SHR
//LKED.SYSLMOD DD SP ACE= (1024, (600,50,1) ,RLSE) ,DISP= (NEW, CATLG) ,
// UNIT=DISK,VOL=SER=XXXXXX,DSN=TSOID#. CANDE. LOAD(CANDE)
//LKED.SYSIN DD *
ENTRY MAIN
INSERT MAIN,PRHERO,RESOUT,PRINC,ESTAB
OVERLAY A
INSERT STEEL, ALUMIN,EMOD
OVERLAY B
INSERT HINGE, SETU,INVER
OVERLAY A
INSERT CONCRE,CONMAT
OVERLAY A
INSERT PLASTI, BASIC
OVERLAY C (REGION)
INSERT PREP
OVERLAY D
INSERT CANB0X,CAN1,XCAN,XCAN2
OVERLAY D
INSERT GENNOD, SAVEG , GENEL , GENEND , AF ,MODMSH , SAVED
OVERLAY C
INSERT HEROI C , CONVT , GEOM , RESP IP , REDUCI , BAKSUB , XFACES
OVERLAY D
INSERT READM,ANISP, DUNCAN, HARD IN, INTP1
INSERT STIFNS , BEAMEL , STFSUB , STRESS , BEAMST
OVERLAY C
INSERT BURNS
/*
Note: Overlay commands start in column 2.
206
TABLE D.4  JCL to run CANDE
//CANDERUN JOB (XX,XXXX, ,10) ,IDNUMBER,NOTIFY=TSOID#,
// REGION=256K,TIME=2
//STEP1 EXEC PGM=CANDE
//STEPLIB DD UNIT=DISK,VOL=SER=XXXXXX,DISP=SHR,
// DSNTS0ID#. CANDE. LOAD
//FTO5F001 DD UNIT=DISK,VOL=SER=XXXXXX,DISP=SHR,
// DSN=TSOID#.PROBNAME.DATA,
// DCB=(RECFM=FB,LRECL=80,BLKSIZE=3120,BUFNO=1)
//FT06F001 DD SYSOUT=A
//FT10F001 DD UNIT=DISK,DSN=&&TEMPO,DISP= (NEW, DELETE),
// SPACE=(TRK, (10,5)) ,DCB=(RECFM=VBS,BLKSIZE=8000,BUFNO=2)
//FT11F001 DD UNIT=DISK,DSN=&&TEMP1,DISP= (NEW, DELETE),
// SPACE=(TRK,(10,5)),DCB=(RECFM=¥BS,BLKSIZE=8000,BUFNO=2)
//FT12F001 DD UNIT=DISK,DSN=&&TEMP2,DISP= (NEW, DELETE) ,
// SPACE=(TRK, (10,5)),DCB=(RECFM=VBS,BLKSIZE=8000,BUFNO=2)
//FT13F001 DD UNIT=DISK,DSN=&&TEMP3,DISP= (NEW, DELETE) ,
// SPACE=(TRK, (10,5)),DCB=(RECFM=VBS,BLKSIZE=8000,BUFNO=2)
//FT30F001 DD UNIT=DISK,DSN=&&TEMP5,DISP= (NEW, DELETE) ,
// SPACE=(TRK, (10,5)),DCB=(RECFM=VBS,BLKSIZE=8000,BUFNO=2)
TABLE D.5  Efficiency comparisons of executing CANDE
Compiler
&
Overlay
Region
size
(K bytes)
Central
Processor
time
(min:sec)
Disk
time
(sec)
Total
cost
(dollars)
G
No overlay
380
1:43
35.0
$9.62
G
Overlay
320
1:44
35.0^
$9.34
HX
No overlay
348
1:03
35.0
$6.45
HX
Overlay
264
1:04
35.0
$6.14
207
J
REFERENCES
1. Katona, M.G., et al., "CANDE  A Modern Approach for the Structural
Design and Analysis of Buried Culverts," Report No. FHWARD775,
Federal Highway Administration, Washington, D.C., October 1976.
2. Katona, M.G. and Smith, J.M., "CANDE User Manual," Report No. FHWA
RD776, Federal Highway Administration, Washington, D.C., October
1976.
3. Katona, M.G. and Smith, J.M., "CANDE System Manual," Report No.
FHWARD777, Federal Highway Administration, Washington, D.C.,
October 1976.
4. Building Code Requirements for Reinforced Concrete, American Concrete
Institute, ACI 31877, 1977, Detroit.
5. Commentary on Building Code Requirements for Reinforced Concrete,
American Concrete Institute, ACI 31877, 1977, Detroit.
6. Vittes, Pedro D., "Finite Element Analysis of Reinforced Concrete
Box Culvert," Master's Thesis, Dept. of Civil Eng. , Univ. of Notre
Dame, May, 1980.
7. Wang, C.K. and Salmon, D.G., Reinforced Concrete Design , Intext
Educational Publishers, Second Edition.
8. Sozen, M.A. and Gamble, W.L., "Strength and Cracking Characteristics
of Beams with #14 and #18 Bars Spliced with Mechanical Splices,"
American Concrete Institute Journal, Detroit, December 1969, pp.
949956.
9. Berwanger, C. , "Effect of Axial Load on the MomentCurvature Relation
ships of Reinforced Concrete Member," SP5011, American Concrete
Institute, Detroit 1975, pp. 263288.
10. Gergely, P. and Lutz, L.A. , "Maximum Crack Width in Reinforced Con
crete Flexural Members," SP20, American Concrete Institute, Detroit
1968, pp. 87117.
11. Lloyd, J. P., Rejali, H.M. and Kesler, C.E., "Crack Control in One
Way Slabs Reinforced with Deformed Welded Wire Fabric," American
Concrete Institute Journal, Detroit, May 1969, pp. 366376.
12. LaTona, R.W. and Heger, F. J. , "Computerized Design of Precast Rein
forced Concrete Box Culverts," Highway Research Record, Number 443,
pp. 4051.
13. Boring, M.R., Heger, F.J. and Bealey, M. , "Test Programs for Evaluating
Design Method and Standard Designs for Precast Concrete Box Culverts
with Welded Wire Fabric Reinforcing," Transportation Research Record
518, pp. 4963.
208
34 • Simpson Gumpertz and Heger Inc., "Report of Test Programs for
Evaluation of Design Method and Standard Designs for Precast
Concrete Box Culverts with Welded Wire Fabric Reinforcing,"
submitted to American Concrete Pipe Association, July 1973.
15. Heger, F.J. and Saba, B.K., "The Structural Behavior of Precast
Concrete Pipe Reinforced with Welded Wire Fabric," Progress
Report No. 2, Project No. 17734, Cambridge, Massachusetts, July
1961.
16. Heger, F.J., "A Theory for the Structural Behavior of Reinforced
Concrete Pipe," Thesis submitted to the Department of Civil Engi
neering, Massachusetts Institute of Technology, January 1962.
17. Heger, F.J., "Structural Behavior of Circular Reinforced Concrete
PipeDevelopment of Theory," Journal of the American Concrete
Institute, November 1963, pp. 15671613.
18. American Society for Testing Materials, "Standard Specification for
Reinforced Concrete Culvert, Storm Drain, and Sewer Pipe," (ASTM
Designation C7670), 1970.
19. Breton, J.M. , "Precast Box Culvert Project  Fabric Materials Test,"
Report to Frank Smith  Gifford Hill Pipe Company, April 1973.
20. American Society for Testing Materials, "Standard Specifications
for Welded Wire Fabric for Concrete Reinforcement," (ASTM Desig
nations A18573), 1973.
21. American Society for Testing Materials, "Standard Specification
for Precast Reinforced Concrete Box Sections for Culverts, Storm
Drains, and Sewers," (ASTM Designation C78976), 1976.
22. American Association of State Highway and Transportation Officials,
"Interim Specification for Precast Reinforced Concrete Box Sections
for Culverts, Storm Drains and Sewers," (AASHTO Designation: M 259
761), 1976.
23. Girdler, H.F. , "Loads on Box Culverts Under High Embankments,"
Research Report 386, Department of Transportation, Division of
Research, Lexington, Kentucky, April 1974.
24. Russ, R.L. , "Loads on Box Culverts under High Embankments: Positive
Projection, without Imperfect Trench," Research Report 431, Depart
ment of Transportation, Division of Research, Lexington, Kentucky,
August 1975.
25. Allen, D.L., and Russ, R.L. , "Loads on Box Culverts under High
Embankments: Analysis and Design Considerations," Research Report
491, Department. of Transportation, Division of Research, Lexington,
Kentucky, January 1978.
209
J
26. Kulhawy, F.H., J.M. Duncan, and H.B. Seed, "Finite Element Analysis
of Stresses and Movements in Embankments during Construction," U.S.
Army Eng. Waterways Experiment Station, Contract Report 5698,
Vicksburg, Miss., 1969.
27. Duncan, J.M. , and C.Y. Chang, "Nonlinear Analysis of Stress and
Strain in Soils, Journal of Soil Mechanics and Foundations Div. ,
ASCE, vol. 96, No. SM5, Sept. 1970, pp. 16291653.
28. Wong, Kai S. and J.M. Duncan, "Hyperbolic StressStrain Parameters
for Nonlinear Finite Element Analysis of Stresses and Movements in
Soil Masses," Report No. TE743, University of California, Berkeley,
July 1974.
29. Duncan, J.M. , et . al. , "Strength, StressStrain and Bulk Modulus
Parameters for Finite Element Analyses of Stresses and Movements in
Soil Masses, Report No. UCB/GT/ 7802 to National Science Foundatidn,
April 1978.
30. Lee, CheeHai, "Evaluation of Duncan's Hyperbolic Soil Model,"
Master's Thesis, University of Notre Dame, May, 1979.
31. Katona, M.G., £t. al . , "Structural Evaluation of New Concepts for
LongSpan Culverts and Culvert Installations," FHWA Report No. RD
79115, Washington, D.C. , December, 1979.
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«U.S. GOVERNMENT PRINTING OFFICE:
FEDERALLY COORDINATED PROGRAM (FCP) OF HIGHWAY
RESEARCH AND DEVELOPMENT
The Offices of Research and Development (R&D) of
the Federal Highway Administration (FHWA) are
responsible for a broad program of staff and contract
research and development and a Federalaid
program, conducted by or through the State highway
transportation agencies, that includes the Highway
Planning and Research (HP&R) program and the
National Cooperative Highway Research Program
(NCHRP) managed by the Transportation Research
Board. The FCP is a carefully selected group of proj
ects that uses research and development resources to
obtain timely solutions to urgent national highway
engineering problems.*
The diagonal double stripe on the cover of this report
represents a highway and is colorcoded to identify
the FCP category that the report falls under. A red
stripe is used for category 1, dark blue for category 2,
light blue for category 3, brown for category 4, gray
for category 5, green for categories 6 and 7, and an
orange stripe identifies category 0.
FCP Category Descriptions
1. Improved Highway Design and Operation
for Safety
Safety R&D addresses problems associated with
the responsibilities of the FHWA under the
Highway Safety Act and includes investigation of
appropriate design standards, roadside hardware,
signing, and physical and scientific data for the
formulation of improved safety regulations.
2. Reduction of Traffic Congestion, and
Improved Operational Efficiency
Traffic R&D is concerned with increasing the
operational efficiency of existing highways by
advancing technology, by improving designs for
existing as well as new facilities, and by balancing
the demandcapacity relationship through traffic
management techniques such as bus and carpool
preferential treatment, motorist information, and
rerouting of traffic.
3. Environmental Considerations in Highway
Design, Location, Construction, and Opera
tion
Environmental R&D is directed toward identify
ing and evaluating highway elements that affect
• The complete sevenvolume official statement of the FCP is available from
the National Technical Information Service, Springfield, Va. 22161. Single
copies of the introductory volume are available without charge from Program
Analysis (HRD3), Offices of Research and Development, Federal Highway
Administration, Washington, D.C. 20590.
the quality of the human environment. The goals
are reduction of adverse highway and traffic
impacts, and protection and enhancement of the
environment.
4. Improved Materials Utilization and
Durability
Materials R&D is concerned with expanding the
knowledge and technology of materials properties,
using available natural materials, improving struc
tural foundation materials, recycling highway
materials, converting industrial wastes into useful
highway products, developing extender or
substitute materials for those in short supply, and
developing more rapid and reliable testing
procedures. The goals are lower highway con
struction costs and extended maintenancefree
operation.
5. Improved Design to Reduce Costs, Extend
Life Expectancy, and Insure Structural
Safety
Structural R&D is concerned with furthering the
latest technological advances in structural and
hydraulic designs, fabrication processes, and
construction techniques to provide safe, efficient
highways at reasonable costs.
6. Improved Technology for Highway
Construction
This category is concerned with the research,
development, and implementation of highway
construction technology to increase productivity,
reduce energy consumption, conserve dwindling
resources, and reduce costs while improving the
quality and methods of construction.
7. Improved Technology for Highway
Maintenance
This category addresses problems in preserving
the Nation's highways and includes activities in
physical maintenance, traffic services, manage
ment, and equipment. The goal is to maximize
operational efficiency and safety to the traveling
public while conserving resources.
0. Other New Studies
This category, not included in the sevenvolume
official statement of the FCP, is concerned with
HP&R and NCHRP studies not specifically related
to FCP projects. These studies involve R&D
support of other FHWA program office research.
DOT LIBRARY
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